Transactions A.S.M.E.

Transactions
of the
A.S.M.E.
Purchase and Use of Fuel E. W. Stone 639
Salt-Velocity Measurements at Low Velocities in Pipes
Field Checks of the Salt-Velocity Method . .
L. J . Hooper 6 51
0. H. Dodkin 663
The Viscosity of Superheated Steam . G. A. Hawkins, H. L. Solberg, and A. A. Potter 6 77
Steam-Turbine Blading R. C. Allen 689
Discussion of Attack on Steel in High-Capacity Boilers as a Result of Overheating Due
to Steam Blanketing E. P. Partridge and R. E. Hall 7 1 1
Determination of the Purity of Steam by Gravimetric and Spectrographic Methods
......................................................M. C. Schwartz, W. B. Gurney, and T. E. Crossan 7 19
Determination of Ammonia in Condensed S t e a m ..........................................................
......................................................M. C. Schwartz, W. B. Gurney, and T. E. Crossan 723
Determination of Purity of Steam by the Electrolytic-Conductivity Method . . . .
......................................................W. B. Gurney, M. C. Schwartz, and T. E. Crossan 728
NOVEM BER, 1940
VOL. 62, NO. 8

Transactions
of The American Society of Mechanical Engineers
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P urchase and Use of Fuel
By E. WADSWORTH STONE,1 THOMPSONVILLE, CONN.
In this paper, the author discusses the location of in ­
dustry, sources of its fuel supply, available com peting
fuels and their relative cost, the purchaser’s com bustion
equipment, the m ethod and efficiency of conversion, all
of which are factors influencing the cost of power and the
fuel necessary to produce it. The history of fuel production,
supply, regulation and use, and the substantial im ­
provement in com bustion equipm ent during the last
twenty years afford an interesting commentary upon the
complicated problem of fuel purchase and supply.
FUEL represents potential energy for release in the form of
heat, and its purchase and use involve many interrelated
factors.
Consumers evaluate fuel according to the recoverable heat
when this is converted in their own equipment and for their own
use. Consumer interest will vary to the extent that certainty of
fuel supply and fuel cost are important factors in their own
production problem.
Producers, to continue in business, must realize in the aggregate
their true actual cost of fuel production, preparation, and
sale. Producers must also accurately appraise markets, consumer
needs, and available conversion equipment in the preparation
and merchandising of their fuels successfully to meet competition.
Consuming markets seldom are located at or adjacent to the
sources of supply. The cost of transportation, definitely established
for whatever means employed, therefore constitutes a
necessary, measurable, and usually a major factor in the delivered
fuel cost.
The cost of production, distribution, and utilization thus determines
the character and extent of the fuel markets, price
levels, and the relative values of competing fuels available in any
common consuming market.
Regulation, either state or federal, provided it properly
recognizes the public, producer, and consumer interests, and fair
existing competitive opportunities, may prove helpful. But
regulation which ignores the effects of changing economic conditions
and disregards the ordinary flow of fuel, in response to
normal supply and demand under normal conditions of free and
open competition, may cause a shift of markets and a permanent
dislocation of production and consumption sufficient to offset any
immediate temporary or localized gain.
Integrated valuations of fuel by producer and by consumer,
measured one against the other, will determine its fair average
value. Any abnormal or arbitrary departure therefrom will
adversely affect the producers of fuel and the consuming markets
they serve.
This country’s growth, commencing at the Atlantic seaboard,
has largely been toward the West and Northwest, with industry’s
movement and location generally determined by the availa-
bility of raw materials, consuming markets, and convenient
means for the distribution of manufactured products.
Fortunately, acceptable fuels were usually accessible. Fortunately
too, excepting a few industries like steel and glass,
fuel constitutes but a minor portion of the raw-material requirement.
Moreover, the development of power represents but a
small, although increasing, part of the cost of production, due
to the continued advance in the cost of manual labor. Transportation
factors governing commodity movement were also
favorable toward the convenient supply of fuels.
Initially, industry, particularly in New England, was located
to make use of available water power. Increasing power requirements
later necessitated the installation of steam relay
stations and, eventually, the use of purchased power.
The gradual movement of industry westward, and its convenient
location along rivers, lakes, railroads, and other available
means of transportation, has brought about a corresponding
redistribution of industrial-fuel markets and, likewise, shifts in
the centers of population and domestic-fuel markets.
In each instance, the continued effort to secure lower costs of
heat and power has stimulated the use of more efficient combustion
equipment and lower grades and less expensive kinds of fuels.
From 1889 to 1929, energy for all uses increased approximately
sixfold; from 1909 to 1929, industrial power requirements increased
some twofold; and in the last 20 years, the use of electric
energy has increased threefold. This increase in the use of power,
particularly of electrical energy, has brought about the installation
of many large generating stations in and near important
industrial centers.
Extensive transmission systems have been provided to interconnect
large steam-generating plants and the near-by hydrostations.
In spite of this augmented flexibility of generation and
supply of electrical energy, the relatively limited storage facilities
at the source and at consumers’ plants require continued and
even greater dependence upon the transportation systems for
high-speed service, reliability, and continuity of fuel supply.
Frequently, reference is made to the location of large generating
stations at the mine or other source of fuel supply, but the fueltransportation
factor may be offset by the lack of condensing
water at the mine and the economics of electrical-energy transmission.
Bituminous coal has been and undoubtedly will continue
to be the principal source of steam and electrical energy.
Competing fuels and hydropower are making substantial increases,
although seasonal variations in the latter may necessitate
steam relay stations. Calculated at the prevailing fuel rate for
electrical energy, the percentage of hydropower to the total
energy consumption has approximated 3 to 4 per cent (1, 3)s
since 1899.
L o c a t io n — I n d u s t r y a n d F u e l S u p p l y
1 Research and Consulting Engineer, Bigelow-Sanford Carpet
Company, Inc. Mem. A.S.M.E.
Contributed by the Fuels Division and presented at the Spring
Meeting, Worcester, Mass., May 1-3, 1940, of T h e A m e r i c a n S o ­
c ie t y o f M e c h a n i c a l E n g i n e e r s .
N o t e : Statements and opinions advanced in papers are to be
understood as individual expressions of their authors, and not those
of the Society.
C o m p e t i n g F u e l s
Coal deposits, with minor exceptions, are conveniently accessible
to all consuming markets. On the other hand, oil is produced
principally in the states of California, Oklahoma, and
Texas, and to a lesser degree in several other states. The producing
areas for natural gas roughly coincide with those for oil.
Bituminous coal, fuel oil, and natural gas not only compete with
each other in those markets easily reached or favorably situated,
but with domestic anthracite in the East and with manufactured
2 Numbers in parentheses refer to the Bibliography at the end of
the paper.

640 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
gas and coke in many industrial centers throughout the United
States.
"New England is a good example of a competitive fuel market.
Imported fuels are a negligible factor, but the production of
hydropower and the use of fuel oil, anthracite, and bituminous
coal, together with the by-product forms of manufactured gas
and coke, present to both the industrial and domestic consumer
a wide variety of fuels from which to choose.
The importation of oil, although frowned upon by producers of
competing fuels, has the effect of preserving our country’s
rather limited oil reserves. On the other hand, the importation
of foreign coal, particularly anthracite along the Atlantic seaboard,
does detract from the potential markets for solid fuel, of
which there is available several thousand years’ supply.
T A B L E 2
E S T IM A T E D A N N U A L N O R M A L PR O D U C T IO N
B itum inous coal......................................................... 440 to 480 million tons
A n thracite.................................................................... 50 to 60 million tons
P etroleum ............................................................... 1280 to 1300 million bbl
N atural gas..................................................................... 2450000 million cu ft
H ydropow er.................................................................... 45000 million kwhr
N o t e : Coke and m anufactured gas involve the consum ption of other
prim ary forms of fuel and are not included.
nage is produced in the eastern or Appalachian fields, 18 per cent
in the middle-western fields (nrinimum price areas (2, 10) Nos. 1
and 2, respectively), and the balance in varied and scattered
amounts in the South and in the Rocky Mountain and northwestern
states.
The annual consumption of bituminous coal in the market
areas east of the Mississippi River, for industrial and domestic
uses, excluding export, railroad, and bunker fuel, approximates
3/i of the total tonnage of the United States, apportioned (4) as
shown in Table 3.
° R efer also to Figs. 1 and 2.
b 26.2 million B tu per to n of bitum inous coal.
The total annual supply of energy from mineral fuels and
water power (at the prevailing central-station heat equivalent)
closely approximates 25,000 trillion Btu. As now proportioned,
this may be expressed in terms given in Table 2.
Approximately 70 per cent of the total bituminous-coal ton-
T A B LE 3 A N N U AL C O N S U M P T IO N OF B IT U M IN O U S COAL IN
EAST
M illion tons
New England S tates......................................................................... 17
New Y ork S ta te .............................................................................#.
Pennsylvania, W est Virginia, and other N orth A tlantio
23
S tates.................................................................................................
O ther states south of th e Ohio and east of th e M ississippi
75
R ivers................................................................................................
S tates north of th e Ohio and east of th e M ississippi Rivers
42
162
Substantially all the anthracite is produced in eastern Pennsylvania
and consumed in the North Atlantic States.
The principal consumption of fuel and gas oil is in the producing
and adjacent areas and those markets conveniently

STONE—PURCHASE AND USE OP FUEL 641
reached via tidewater and pipe line. The total distribution (5)
in 1937 approximated:
Per cent
New England S tates............................... 9
Middle Atlantic S tates........................... 29
South A tlantic S tates............................. 4
N orth C entral S tates.............................. 14
South Central S tates.............................. 20
Rocky M ountain S tates......................... 1
Pacific C oast S tates................................ 23
T o tal...................................... 100
The consumption of natural gas likewise is generally confined
to the producing areas, except where accessible by pipe
line, with the total distribution (5) 1937 divided as follows:
Per cent
California................................................... 14
Louisiana.................................................... 13
Oklahoma................................................... 12
Texas........................................................... 36
W est Virginia............................................ 6
Other states............................................... 19
T o tal...................................... 100
In 1937, there was produced from fuels 77,348 million kwhr
of electrical energy, and from water power, in competition with
fuels, some 43,702 million kwhr, a total production of 121,050
million kwhr,3 apportioned (6) as shown in Table 4.
* Approxim ately 128,000 million kwhr in 1939.
F i g . 2
R e l a t iv e R a t e o f G r o w t h o f A n t h r a c it e , B it u m in o u s C o a l, P e t r o l e u m , N a t u r a l G a s , a n d W a t e r P o w e r i n
t h e U n it e d St a t e s
(Index num bers 1918 equals 100. D a ta taken from supplem ent to W eekly C oal R eport No. 1084 of T he N ational B itum inous Coal Commission.)
N otes on F ig . 2 ^ U nit h eat values used
* C.R . W ater power a t constant ra te ;---------------------- K ind of fuel B tu value
v 4 02 lb per kw hr
iiS o W ater Power a t prevailing ra te ;----------------------
A nth racite.................. ....................................................
B itum inous coal............................................................
13600 per lb
13100 per lb
te ’ ,3 ®6 *b Per,
1937 rate, 1.42 lb per kw hr
P etroleum .......................................................................
N atural gas.....................................................................
6000000 per bbl
1075 per ou ft
W ater power equivalent (C.R.) is based on th e year 1913 or m id-point of years compiled (1889-1937) an d is the constant factor of 4.02 lb of coal per kw hr
W ater power equivalent (P.R .) is based on the prevailing factor which varies from 7.05 lb of coal per kw hr in 1899 to 1.42 lb of coal per kw hr in 1937
Wnen production of w ater power was not available, th e horsepower equivalent was used, taken a t 20 per cent capacity factor for m anufacturing and mines
and 40 per cent for utilities

642 TRANSACTIONS OF TH E A.S.M.E. NOVEMBER, 1940
T A B L E 4
A P P O R T IO N M E N T O F E L E C T R IC A L -E N E R G Y
P R O D U C T IO N IN 1937
H ydroplants, Fuel plants, T otal,
per cent per cent per cent
New England S ta te s..................................... 7 6 7
M iddle A tlantic S ta te s.................................
E a st N orth C entral S ta te s..........................
19
5
29
34
26
23
W est N orth C entral S ta te s......................... 5 7 6
South A tlantic S ta te s................................... 16 9 12
E ast South C entral S ta te s.......................... 8 2 4
W est S outh C entral S ta te s......................... 1 8 5
R ocky M ountain S ta te s............................... 10 2 5
Pacific C oast S ta te s................................ 29 3 12
T o tals......................................................... 100 100 100
For comparison, it is interesting to note that the electricgenerating
capacity of utility and other power and light plants
devoted to public use at the end of the year 1937 was reported to
be as follows:
10.474.000 kw for hydroplants
26.558.000 kw for those using fuel
37.032.000 kw4 for entire United States
The total amount and that for the two methods of generation are
apportioned (6), as shown in Table 5.
TABLE 5 APPORTIONMENT OF ELECTRIC-GENERATING
CAPACITY IN 1937
H ydroplants, Fuel plants, T otal,
per cent per cent per cent
New E ngland S ta te s..................................... 8 8 8
M iddle A tlantic S ta te s................................. 14 29 25
E a st N orth C entral S ta te s......................... 6 28 22
W est N orth C entral S ta te s......................... 5 9 8
South A tlantic S tates.................................... 19 9 12
E ast South C entral S ta te s.......................... 10 3 5
W est South C entral S ta te s......................... 1 6 5
R ocky M ountain S ta te s............................... 11 2 4
Pacific C oast S ta te s ....................................... 26 6 11
T o tals......................................................... 100 100 100
Those portions of the fuels which compete with each other and
with hydropower may be conveniently expressed in terms of
“equivalent bituminous-coal tonnage”6 for a normal year.
These values are given in Table 6.
TABLE 6 “EQUIVALENT BITUMINOUS-COAL TONNAGE” FOR
COMPETITIVE POWER SOURCES IN NORMAL YEAR
Million tons
Bituminous coal:
Domestic.......................................................... 88 to 96
Railroad............................................................ 88 to 96
Industrial and power use.................................. 264 to 288 440 to 480
Anthracite:
Domestic.......................................................... 33 to 39
Industrial.......................................................... 17 to 21 50 to 60
Fuel and gas oil:
Domestic....................................................................... ...26
Power............................................................................ 26
Railroad......................................................................... 16
Other............................................................................ 35 103
Natural gas:
Domestic....................................................................... 20
Industrial...................................................................... 39 59
Hydropower............................................................................. 31
683 to 733
E fficiency of Conversion
The improvement in efficiency of converting the potential
energy of coal and other fuels into electrical, mechanical, or other
useful forms of energy offers an interesting commentary upon the
relatively constant total supply of energy from mineral fuels
since 1918, except during the depression years of the 1920’s and
again in the 1930’s. The following examples will illustrate this
point:
The average number of pounds of coal required per 1000 gross
ton-miles for freight service on steam railroads decreased from an
4 Approximately 40 million kw at the end of 1939.
s 26,200,000 Btu per net ton.
average of 170 lb in 1919-1920 to 115* in 1938; a reduction of
32 per cent (4, 5, 7).
During the same period, the pounds per passenger-train carmile
decreased from 18.5 lb to 14.9 ;7 an improvement of 19Vj
per cent (4,5,7).
■In 1918, the production of a gross ton of pig iron required the
use of the equivalent of some 3577 lb of coal, whereas, in 1938
this amounted to but 2865 lb; an improvement of nearly 20
per cent (4,5).
The improvement in the efficiency of production of electrical
energy is even more marked. It is reported that in 1889 the
average consumption of coal was 7.5 lb per kwhr. This requirement
(1, 4, 5, 6, 8, 11) was reduced in
1902 to 6.4 lb per kwhr
1912 to 4.5 lb per kwhr
1919 to 3.22 lb per kwhr
1920 to 3.04 lb per kwhr
1929 to 1.69 lb per kwhr
1938 to 1.41 lb per kwhr
The reduction since 1919 approximates 56 per cent.
Several central stations are now producing electrical energy for
0.75 lb of coal per kwhr, approximately 1/ 2 the 1938 average for
the entire United States. Many factors have encouraged this
improvement in efficiency and undoubtedly will continue to do
so, although perhaps not to the extent possible in the past.
E q u ip m e n t a n d C h a n g e s
The record is replete with coal analyses, boiler- and powerequipment
tests, specifications and descriptions in general and in
detail, including design, construction, and operation of modern
steam-generating stations for the utilization of solid, liquid, or
gaseous fuels, singly and in combination, and need not be further
amplified in this paper.
Considerable thought has been given to the storage of fuel,
its delivery to the furnace, the method of firing, combustion control,
the removal of waste, desirable fuel characteristics, boiler
design, operating pressures, capacities, and the rate of heat release.
Much study has been devoted to size and arrangement of
combustion space; gas velocities, to secure maximum rate of
heat transfer; furnace temperatures; the use of waterwalls and
special refractories to withstand the higher prevailing furnace
temperatures; the use of alloy steels and welded construction;
pulverizer design and the use of watercooled stokers.
Heretofore, the quantity of ash and of sulphur and the ashfusion
temperatures appeared to be limiting factors in the use of
certain types of combustion equipment. Methods for continuous
or intermittent ash removal, either in the dry or plastic state,
have been improved. We know more about what actually takes
place in the combustion of fuel and the behavior of ash on refractory
and boiler surfaces. This has indicated that the slagging
characteristics of a coal may be dependent upon the fusion temperatures
of its component parts rather than an average of all,
and that Fe20 8, lime, and other components of ash definitely influence
slag formation. All this has helped to make possible the
combustion of high ash pulverized coal with ash-softening
temperatures as low as 2300 F, in a dry-bottom furnace, and,
conversely, the use of high-fusion coals with ash-softening temperatures
as high as 2550 F, in wet-bottom furnaces, the furnace
temperatures in either case being controlling factors.
These numerous experiments, studies, and investigations, individually
and collectively, constitute an important contribution to
the use, operation, and efficiency of modern steam-generation
equipment, and encourage still further improvement. The ex-
« 112 in 1939.
» 14.8 in 1939.

STONE—PURCHASE AND USE OF FUEL 643
perience and results obtained on the larger units have been and
can be applied, with perhaps less refinement, to the smaller and
simpler combustion units for industrial use, or for the modification
and more efficient utilization of existing equipment.
Large utility and industrial consumers of fuel employ technicians
to check fuels and power-plant operation carefully and
frequently. To the small consumer, who may question the
economic justification of such a check, there are now available
fuel technicians and an extensive array of operating data. The
larger producers of fuel maintain, for the convenience of their
customers, technical staffs thoroughly acquainted with their fuels
and frequently experienced in actual plant operation. The
National Association of Purchasing Agents (9) has also reminded
its membership of the more important factors influencing
the use of fuels.
Heretofore, initial cost has frequently been permitted to influence
the installation and use of steam-generating equipment,
often to the extent of restricting it to a single or a limited number
of types or grades or sizes of fuel.
There appears to be a growing realization among equipment
manufacturers and, likewise, among the users of all kinds of fuel
that, with very little additional effort and expense, steamgenerating
equipment, having a much wider range of fuel utilization,
can now be manufactured and installed.
If this be a correct interpretation of the trend, this increase in
the adaptability and range of use of fuel-burning equipment will
enable the consumer, in spite of the tendency to regulate both
price and source of supply, to purchase fuel at a low average unit
cost and effect the same or even greater efficiencies of conversion
than have heretofore prevailed.
Likewise, manufacturers may, at little added cost, modify their
combustion equipment to make possible the burning of solid,
liquid, or gaseous fuel, and the high- and low-grade fuels of any
type with substantially equal efficiency.
Thus, the consumer will be enabled to take greater advantage
of the potential and changing fuel markets and available methods
of transportation, and in his purchases can be governed to a
larger degree by the actual fuel value measured by the delivered
fuel cost per million Btu.
M a r k e t T r e n d s 8
Natural gas has been marketed in ever-increasing quantity
since 1921, similarly fuel oil, with the increased consumption
more pronounced for domestic heating than for other uses. At
the same time, there has been a definite trend downward in the
use of both anthracite and bituminous coal, this trend being morfe
pronounced since 1926.
Throughout the last 20 years, steam-generated electrical
energy has increased threefold, with but slight increase in the
use of fuel, due to improved efficiency in fuel utilization. Hydrogenerated
electrical energy has increased in substantially the
same proportion, but, if evaluated at the prevailing rate for
steam-generated energy, the coal displacement appears to show
relatively small change since 1918.
Except for the periods during and immediately following the
World War, and during strikes and other abnormal times, all
bituminous coal has averaged to realize at the mine between
$1.75 and $2 (5) per net ton.
The realization at the mine for Pennsylvania anthracite appears
to have declined steadily from an average of 15.35 in 1931
(5) to nearly $4 in 1937. The downward change, which was
greater for domestic sizes, was somewhat offset by a gradual increase
in the realization for industrial and steam sizes.
If we take the tidewater price of bunker C fuel oil at New
* Refer to Figs. 1 and 2.
York, the average price per barrel for the years 1933 to 1939,
inclusive, has ranged from approximately $1 to $1.30 (5), with
the actual price, of course, depending somewhat upon the local
market conditions, the quantities involved, and subject to
greater fluctuation than that of other fuels.
The average annual realization for natural gas, on the other
hand, has ranged within relatively narrow limits, from approximately
211/ 2 cents to 243/ 4 cents per thousand cubic feet at point
of consumption for the years 1925 to 1938 (5), inclusive. Although
the figure for all natural gas has remained substantially
constant, the average for industrial use during the same period
appears to have declined from a high of nearly 13 cents to a low
of 9.7 cents, whereas, the corresponding rates for domestic consumption
appear to have increased from a low of 56 cents to a
high of some 69 cents per thousand cubic feet.
These figures, unless otherwise stated, are merely averages for
the entire United States. They do not indicate the relative
prices at any particular market or destination, where the different
transportation costs will substantially modify the actual delivered
prices for the available fuels.
The better- and not the lower-grade fuels move to the more
distant markets, where transportation is important. From time
to time, strikes, changing business conditions, temporary shortages
of supply, and other varying factors may serve to modify
consumer preference and the relative prices of competing fuels,
but the general parallelism between the costs of the different types
of fuel (solid, liquid, and gaseous) appears to continue.
Recently, producers of bituminous coal have given more
attention to the screening, sizing, cleaning, and possible dust
treatment of their coals, the better to meet changes in market
conditions and in combustion equipment, realizing that thereby
they can substantially improve their opportunities to hold and
possibly extend existing markets.
The use of the finer sizes of slack and double-screened coals,
respectively, for the industrial and domestic trade has substantially
increased. Slack coals, formerly a drug on the market,
selling at ridiculously low prices, have encouraged the introduction
of pulverized-fuel equipment. The introduction of
domestic stokers has made possible and convenient the use of
smaller and less expensive double-screened sizes of coal, tending
to reduce the supply of slacks available for industrial use.
The price differentials between the so-called “prepared” or
double-screened coals, on the one hand, and slack coals, on the
other, tend to diminish, with the mean or weighted average approaching
but not reaching the mine-run figure. This diminishing
differential as to size, coupled with corresponding differentials
as to grade, has accelerated the price leveling-out process,
as indicated by the generally greater increase proposed under
the new price schedules for slacks and industrial coals than for
the large double-screened and domestic sizes.
T r a n s p o r t a t io n
Freight rates are not always proportional to the distance
traveled, in fact those for long hauls often are relatively lower, to
enable newer and more distantly located coal fields to compete in
the larger consuming markets with the older near-by fields.
The cost of transporting a ton of coal via rail averaged; for
the entire United States, an amount equal to or greater than the
average cost of the fuel itself. This cost ranged from $2.15 to
$2.27 per net ton during the years 1929 to 1938, inclusive, with the
actual transportation charge to many consuming markets several
times the actual fuel cost f.o.b. cars at the mine. Approximately
24 per cent of the bituminous-coal production is water-borne
wholly or partly via lake, river, or tidewater, and 7 per cent is
moved by truck from the mine to near-by destinations and usually
at less than rail rates.

644 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
Transportation cost, reliability of the carrier or carrier system,
and the continuity of supply are important factors when coal or
fuel oil are considered.
The increasing availability of natural gas has encouraged the
extension of pipe lines so that now it may be so conveyed thousands
of miles from the source of supply to the ultimate consumer.
Anthracite destined for eastern markets generally moves via
rail, or rail and tidewater, with a small portion moving via truck.
S e l e c t i o n a n d E v a l u a t io n o p F u e l
Purchasers of fuel are anxious to familiarize themselves with
the limitations of their own equipment, the operating man’s viewpoint,
and the characteristics of available fuels. They desire some
readily understood measure of fuel evaluation and performance,
first, to guide them in its selection and, subsequently, to check
actual performance when the fuel is burned in their own plants.
A suitable measure of performance will facilitate comparison
between plants and between competing fuels in the same plant,
also the effect of variations in preparation and physical characteristics
upon the efficiency of utilization.
A vast amount of detailed information and individual plantoperating
experience has been accumulated and made available
by the engineering profession and by producer and consumer
organizations as to available fuels. Nevertheless, the major
problem still remains.
A scientific approach to the purchase of fuel is possible. Frequently,
however, the power or fuel portion of the total manufacturing
cost is not sufficiently large to justify the extensive and
detailed evaluation methods used by those who consume millions
of tons annually. However, the many chemical and physical
characteristics of fuels undoubtedly do have a marked bearing
upon the selection of appropriate fuels and, having once determined
the desired type or types, the field is considerably narrowed.
It then becomes a m atter of evaluating fuels on the basis
T A B L E 7
SU G G E S T E D M E T H O D O F E V A L U A T IN G FU E L S
of actual heat recovery rather than on a casual analysis of a
salesman.
Coal is now available in a greater variety of sizes and grades
than is generally realized. The producers themselves have
learned to better prepare, size, grade, and treat their coals;
fuel oils have increased in specific gravity and heat content;
refinery by-products and residues have become available; and
the extension of natural-gas lines has brought this fuel into
many otherwise inaccessible markets. Combustion equipment
has been designed or modified to permit high rates of heat release
and to burn not only one but several fuels with substantially
equal facility. The fuel-selection problem thus has become increasingly
complex.
The average consumer is manifestly somewhat bewildered by
the kaleidoscopic changes in fuel prices, fuel markets, and factors
affecting fuel production, distribution, and use. His interest in
the evaluation of fuels leans toward a definite, tangible yardstick
or measure of evaluation and a definite departure from the old
rule-of-thumb method of purchase from “Smith” or “Brown,”
because the fireman thought it burned better or merely because
the supplier had convinced the fireman that his coal shoveled
better than that of his competitor.
The suggested method of evaluation is illustrated in Table 7
for coal, gas, and oil, and for stoker- and pulverized-fuel-fired
boilers.
The variable factors, production, transportation, handling,
and utilization, may be divided into two fundamental and
essential groups as follows:
1 The complete cost of the fuel, including transportation and
conversion, per ton, barrel, or million cubic feet.
2 The heat recovery expressed in convenient units, as, for
example, million Btu per fuel unit.
Both of these in the average industrial plant may be readily
and accurately determined. When compared, they provide a
simple and substantially uniform and sufficiently accurate
method of evaluation regardless of the fuel, the conversion equipment,
the method of operation, plant location, or any other of the
many factors which enter into or effect the utilization of fuel.
After the primary determinations, supplemented by a heat
balance if desired, fuel evaluations are completed for repreeentar
tive fuels and the usual boiler ratings. Subsequent modification
may be quickly approximated to determine the effect of changes in
prices, transportation costs, and operating efficiencies. Thereafter,
periodic tests and an analysis of the results over extended
periods may suggest further changes in fuels and conversion
equipment and make possible the widening of the range of fuel
application. Thus, the management and operating staff are
provided with a yardstick or standard of performance which can
be easily and readily applied to a large or a small plant, and
which includes all the essential fundamental factors. To illustrate,
if we add the following items:
A The delivered cost of fuel, including storage and handling
if necessary;
B Cost of handling the fuel into and through the conversion
equipment;
C Ash- and waste-removal cost;
D Supervision, labor, maintenance, and other operating
expense;
then we can readily determine, from these known and easily
secured data, a unit cost for fuel delivered and consumed.
Actual heat recovery may be readily ascertained from the
steam produced or an equivalent-use evaluation in other processes.
The two predominating and major factors may then be

STONE—PURCHASE AND USE OF FUEL 645
compared directly to determine the cost per million Btu actually
recovered, i.e.
Fuel cost
-------- ---------- = Cost per million Btu recovered
Heat recovered
The three cost factors B, C, and D, are the average for the
particular plant for the preceding year, weighted (plus or minus)
to approximate empirically the probable increase "or decrease, as
the case may be, for handling the particular type, grade, or size of
the fuel in question. The average cost assigned to B, C, and D is
controlling and is a definite and known amount for any given
plant. Such slight modification as may be required properly to
weight these items for any particular fuel is a very small portion
of the total cost of the delivered and converted fuel, in fact so
small as not to be seriously questioned by either the seller or the
consumer of the fuel, and usually small enough to be waived entirely
in case of serious argument.
This method of evaluation includes the cumulative effect of
many factors not otherwise distinguishable except through detailed
analysis. Actual experience and use have shown that it
affords a reasonably simple and accurate comparison of the
energy-conversion cost for different plants, either large or small,
or different usable fuels in the same plant, or even plants other
than those used for the generation of steam.
To a particular producer 5 cents per bbl or 10 cents per ton
differential in the fuel cost (f.o.b. source) may appear to be a
substantial one, and may even mean the loss or gain of his market
or customer. To the consumer, transportation is an added and
an important element of expense; in fact, frequently, as in
New England, transportation cost may average to be twice that
of the fuel itself. Thus, small variations in fuel price, although
seemingly of major importance to the producer, actually become
a minor factor in the consumer’s total cost, and particularly so
if, for the fuels or the plants being compared, there is a marked
difference in the percentage of heat actually recovered.
M in im u m P r i c e s F o r B it u m in o u s C o a l
There has been in progress during the last two years a series
of investigations, collection of fundamental data, and hearings
for the purpose of determining the weighted-average cost of
production and establishing minimum prices for bituminous coal
in accordance with the Bituminous Coal Act of 1937 (2). The
Act authorizes separation of the coal fields of the United States
into 23 producing districts and the grouping of these into 10
minimum-price areas. The Act requires that the production of
Producers’ D a ta to D istrict S tatistical B ureau
(1) Cost of 'production: (1) D irect and indirect expenses plus cost of (16) Coordinated m inim um prices, coal classifications, marketing rules. (16).
selling.
and
to . (17) Proposed as result of coordination by th e several producing districts
for each kind, quality, and size shipped by any of them , in
(4) Preparation: (2) Kinds, qualities, sizes.
Sales: (3) Selling price, spot orders, invoices, credit memo.
com petition into an y and every m arket area, together w ith coordinated
m arketing rules and regulations.
Distribution: (4) Tonnage by kinds, sizes, grades, and use to each
m arket area.
In th e event of failure of th e districts to agree in whole or in part,
(5) District Statistical Bureau to Division Statistical Section:
th e B itum inous C oal Division m ay com plete and propose such
to
coordinated m inim um prices and m arketing rules (17).
(9) Sum m aries of 1936 cost (5), preparation (6), sales (7), and distri- (18) 4~II (b) Hearing:
bution (8) d a ta for each D istrict to statistical section of B itum inous
R egarding inter-district coordination, prices, and m arketing rules
Coal Division; D istrict cost sum m ary (9) to D istrict Boards.
proposed for th e shipm ent of coals into common consuming m arkets.
(10) Adjustm ent 1986 production cost: (19) M arketing rules and regulations: ) P rom ulgated by Secretary of the
D istrict B oards’ recom m endations to D ivision to reflect changes in (20) Prices— m inim um and if necessary,
m axim um :
upon recom m endation of th e
D epartm ent of th e Interior
wages, etc., since 1936.
(11) Cost hearing: (21) D iscounts to distributors
D irector of th e B itum inous
R egrading determ ination weighted average cost of production for
) Coal Division.
each D istrict and price area. (22) Price modifications and changes found necessary or advisable on the
(12) Approved weighted average cost of production authorized by D ivision as and D ivision’s own representations or on request from producer (22) or
basis for proposing m inim um prices for various kinds, qualities, and (23) consum er (23) interests, and after suitable hearing, legal procedure,
sizes of coal produced in each district.
etc.
(13) Intra-district quality and, price relationships to reflect values of th e ’ (24) Records are available to C onsum ers’ Counsel on request.
kinds, qualities, and sizes of coal produced by each mine w ithin th e (25) D istribution or sales agency, registrations.
district as proposed by each D istrict Board, together w ith their (26) Rules: Procedure and relationship betw een D istrict B oard and
(14)
proposals as to m arketing rules and regulations for th e sale of coal.
4-11 (o) Hearing:
and
(27)
Regarding proposed in tra-d istrict q uality and price relationships (28)
and m arketing rules and regulations.
(15) Approved intra-district classifications and marketing rules and regulations.
producers (26) and necessary assistance and inform ation (27).
Tax:
N one on export shipm ents or those to federal, state, or m unicipal
purchasers; 1 cent per ton on all other except th a t produced by
noncode m em bers, which shall be 19*72 cents per ton.

646 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
any district shall realize, as nearly as may be, the weightedaverage
cost of production of the minimum-price area of which
the district forms a part. It also provides for the establishment
of Consumers’ Counsel as a liaison agency between the Coal
Division on the one hand and the consuming public on the other.
The Bituminous Coal Division of the Department of the
Interior on July 1,1939, took over the work of the former National
Bituminous-Coal Commission, and since that time has carefully
and painstakingly continued the procedure (Fig. 3) of establishing
minimum prices, f.o.b. mine, for bituminous coal from all producing
districts within the United States and for movement into the
many consuming market areas (some 200) within its boundaries.
The task proved to be a formidable one; in fact, the final minimum-price
hearing, which in a measure comprised an open
coordination meeting, has taken 6 months, with the accumulation
of some 25,000 pages of testimony and over 1700 exhibits, many
of which are replete with data of varied and complicated character.
All of this, together with the fundamental data upon which the
weighted-average-cost determinations have been predicated,
comprises a more extensive accumulation of pertinent data with
reference to the production, distribution, consumption, and values
of bituminous coal than any heretofore assembled in the history
of the industry, and of inestimable value to producers and consumers
alike.
The magnitude of the problem and the difficulty of arriving at
an equitable solution becomes all the more apparent when one
realizes that it involves:
1 The determination of the cost of producing and selling coal.
2 The classification of all coals by size and by grade.
3 The cost of transportation.
4 The establishment of minimum prices for the various kinds,
qualities, and sizes of coal, produced in the several districts, to
reflect, as nearly as possible, their relative market values at
points of delivery in each common consuming-market area.
In addition, the Bituminous Coal Division, in the absence of
complete agreement between producing districts, found it
necessary to undertake the work of final price coordination.
Production costs for all mines for the year 1936, adjusted to
reflect subsequent changes in wages, etc., including a reasonable
cost of selling, when summarized, were used to determine the
weighted-average cost of production per net ton for each price
area for all of the ascertainable tonnage produced therein.
Subject to public hearing and subsequent approval by the
Coal Commission, each district board proposed a classification of
its coals and differentials as to size and grade. Subsequently, as
a part of coordination, the Coal Division determined the base
coals of each producing district and related these to the base coals
of the other producing districts in their common consuming
markets. The remaining coals of each district were then differentially
related to its own base coals within the same market.
Throughout, consideration was given to sizing and preparation
of coal, methods of cleaning and treatment, coal analyses, and to
the possibility of rearranging and regrouping sizes proposed for
each producing district, to facilitate the relation of various coals,

STONE—PURCHASE AND USE OF FUEL 647
and simplify price structure for both producer and consumer.
Transportation tariffs are intricate in relationship and infinite
in variety. Transportation cost constitutes a very definite and
important factor in determining the f.o.b. mine prices of fuel.
This is necessary if coals from the various sources are to compete
with other fuels on a basis which will insure fair opportunities for
the consumer to purchase and producers to distribute their coal
in their natural markets without serious interruption or dislocation
of tonnage.
In relating truly competitive coals, substantial freight differentials
frequently were reduced by the absorption of a limited
portion of such differential by the expedient of lowering the f.o.b.
mine price of the coal or coals moving on the long freight rates.
Coal for truck delivery was priced at the mine to effect the same
delivered price as that for rail movement, an average trucking
charge being used for this purpose.
An effort has been made to retain the advantages of low-cost
water transportation in some market areas and for those consumers
who are privileged to receive fuel on an “alongside” basis.
Coal moving ex-dock, ex-river, or ex-lake has been coordinated
with competing fuels moving all-rail into the same common consuming
market.
The level of prices in any common consuming market is to a
large degree determined by that market, its proximity to the
producing field, and the extent to which competing fuels, whether
they be gas, oil, other coals, or even hydropower, may affect the
movement of bituminous coal into that market.
Fig. 4 is a reproduction of an exhibit (4) introduced by the Coal
Division during the final price hearing. On this the Division has
noted the principal markets, producing districts, and f.o.b. mine
prices for competitive “base” steam coals for “certain selected
destinations.” This is included here by way of illustration,
without in any way implying approval. I t seemingly indicates
an attem pt so to coordinate coal prices as to avoid discrimination
between classes of consumers, market areas, and industrially
competitive sections of the eastern United States.
The proposed prices involved consideration of both levels and
differentials, and include all sizes and grades of coals, whether
borne by rail, water, or truck, and for the following uses:
a Domestic
b Industrial
c By-product-coke manufacture
d Water-gas and producer-gas manufacture
e Railroad-locomotive fuel
f Vessel and bunker fuel.
The “minimum-price” fixing procedure has reached a point
where the Bituminous Coal Division has been able to estimate the
realization anticipated from the proposed minimum prices, based
on the 1937 tonnage distribution. In some instances, proposed
prices are lower and in others considerably higher than the present
market, with the average realization exceeding that now prevailing.
The anticipated realization, together with the corresponding
weighted-average cost for each of the several producing
districts and price areas, is included in Table 8 (10).
The weighted-average cost of production and corresponding
realization for minimum-price areas Nos. 1 and 2 (comprising the
Appalachian and middle-western producing fields, respectively),
allowing for subsequent modifications suggested by the Marketing
Division, are:
Under free and open competition, relative market values in
each market area involved price differentials as to sizes and
qualities for the coals of each producing district and the level of
prices for that market.
Differentials in price are influenced by quality, size, size consist,
treatment, use, chemical and physical structure, and the
adaptability of a coal or coals to use, markets, transportation, and
other factors.

648 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
within each of the two price areas where they compete in common
consuming markets. This may be illustrated diagrammatically
by Fig. 5, in which is shown the relative weighted-average production
costs and anticipated realizations for price areas Nos. 1
and 2. It so happens that a substantial portion of the tonnage
of the southern districts of price area No. 1 moves into the
Middle West in competition with the coals of price area No. 2,
produced within that consuming area, price area No. 1 coals in
both domestic and steam sizes generally being superior in value.
If the relationships between the coals of price areas Nos. 1 and
2, as determined by the Division, are correct, it is obvious that the
truly competitive coals from each area must be priced f.o.b. the
mine to deliver at their proper relative market value. Such
coordination may be illustrated by the indicated rotation of the
desired price level or realization lines. The rotation of price
area No. 1 line indicates a price decrease for its westbound coals
and a corresponding increase of the remaining coals moving
eastbound since, only in this manner or by a decrease in realization
(Ri), can coordination of price area No. 1 and price area
No. 2 coals be effected.
The higher level of prices of those coals of price area No. 2
which actually compete with the westbound coals of price area
No. 1 further assist this coordination. However, it is obvious
that, for each ton priced above cost, there is a corresponding ton
priced below cost in order that price area No. 2 realization may be
accomplished.
Actual anticipated realizations indicate that price area No. 1
will underrealize and price area No. 2 overrealize its cost. Therefore,
if the coals of these two price areas have been properly related,
any further narrowing of the delivered price differentials
between their competing coals means an abnormal increase in
price area No. 1 coals moving eastbound or an abnormal lowering
of prices for the coals of price, area No. 2, which are not truly
competitive with those of price area No. 1.
Price area No. 2 values cannot approach a point where they are
considered “dump” prices, and the eastbound coals of price area
No. 1, in turn, cannot be substantially raised without at once
throwing an additional abnormal burden upon eastern consuming
markets. This would also result in further discrimination,
as between the market areas involved and a transfer of the
burden of realization to an even greater degree than is appropriate
for those coals, principally slacks, moving into market areas
comprising New England and- the eastern seaboard, which have
no alternative but to take the coals of price area No. 1 or turn to
competing fuels or forms of energy, such as hydropower, oil, and,
in some cases, natural gas. Such competition has already resulted
in a substantial increase in the utilization of oil and gas,
and a decreased use of. bituminous coal and anthracite through-:
out the eastern markets. ."
If the relationships between the' coals of price areas Nos. 1
and 2, as determined by the Division* are correct, and if a narrow^
ing of the differentials betweeni the competitive coals of these
two price areas is attempted in their common consuming markets
either by lowering the level of realization (Ri) of price area No. 1
or by raising the level of realization (fl2) of price area No. 2, then
the underrealization of price area No. 1 and the overrealization
of price area No. 2, with respect to their costs of production, are
aggravated to a still greater degree. '
Distribution figures indicate and market history confirms the
movement, for steam raising or by-product purposes in competition
with other price area No. 1 coals, of the bituminous coals
of eastern Pennsylvania (district No. 1), western Pennsylvania
(district No. 2), northern West Virginia (district No. 3), and the
southern low-volatile coals (district No. 7) in substantial quantities
into New England. New York State, and other portions of
the Atlantic seaboard as far south as Baltimore and Washington.
The southern high-volatile coals of district No. 8 and the lowvolatile
coals of district No. 7, particularly in the double-screened
or prepared sizes, move westward through the Great Lakes and
the northwest territory for “alongside” or “ex-dock” delivery
in competition with the coals from Indiana, Illinois, western
Kentucky, and Iowa (districts Nos. 9,10,11, and 12, respectively,
comprising minimum price area No. 2). Slack coals from districts
Nos. 7 and 8 also move into the industrial markets of Ohio.
C o s t p e r M i l l io n B t u
If the average analyses (4) for the several sizes and grades of
coal be applied to the delivered prices corresponding to the
minimum f.o.b. mine prices proposed by the Bituminous Coal
Division, the fuel cost per million Btu delivered is found to range
from 9 to 14 cents for the mine run, and 9 to approximately
13*/s cents for the slack sizes for those destinations and market
areas located at or close to the producing districts of the Appalachian
coal fields. The delivered slack cost for more distantly
located markets may approach 20 cents per million Btu in the
eastern markets, 21 cents in the South, 25 and even 30 cents in
the Northwest for rail movement, and 15 cents for Great Lakes
destinations and for water-borne fuel. The mine-run figure
averages 1 to 5 cents higher than for slacks. In any market,
the differential in unit delivered cost between the high- and the
low-grade fuels is surprisingly small.
Transportation appears to affect the unit cost for steam coals
in eastern markets to a greater degree than in southern markets.
In the Middle West, rail-transportation charges, both short and
cross-haul, and those for water-borne fuel, together with the
physical characteristics of the fuel and its adaptability for longhaul
movement, evidently combine to make for wide variations
in cost per million Btu for apparently small change in actual heat
content of the fuel.
A comparison of solid, liquid, and gaseous fuels in those
markets where they are actually available and competitive
indicates a delivered cost per million Btu of substantially like
amounts during periods of free and open competition.
G o v e r n m e n t R e g u l a t io n
The World-War period, with its abnormal demand for coal for
the manufacture of steel and munitions, and for industry, fostered
the opening of many mines and thereby increased the accessible
supply. The increased fuel prices prevailing during that period
encouraged more efficient utilization and the use of lower-grade
and cheaper fuels. These trends continued after the war so
that the large potential supply of fuel became a serious burden;
the solid-fuel industry, particularly bituminous coal, became
demoralized, and the less efficient and higher-cost mines were
forced out of business. The closing of mines continued, but not
rapidly enough to prevent further demoralization of the fuel
industry. Competition was keen and prices were low. These
conditions, together with the realization that (a) the more accessible
eastern coal reserves faced earlier depletion; (6) actual
fuel recovery warranted substantial improvement; (c) improved
working conditions and wages were in order throughout the
industry, initiated and encouraged governmental interest (11).
This governmental interest continued from the time of the
Hammond Commission in 1922 to, and culminated in, the NRA
regulated prices. These were established at a sufficiently high
level to permit increased wages in the northern and southern
fields together with a general unionization of both. The effort to
stabilize conditions within the industry, then begun, has continued
until the present time. There is now imminent Federal
regulation of the industry through the establishment of minimum
prices for bituminous coal throughout the United States as
prescribed in the Coal Act of 1937 (2).

A recognition of the rather limited petroleum reserves, overproduction
of crude oil, and other economic factors has also
aroused considerable interest on the part of governmental
agencies in an effort to regulate (12, 13) the supply and stabilize
prices for both crude and fuel oil and for oil products.
For similar reasons, attention has been directed to the possible
control of natural-gas pipe-line installations and a stabilization
and regulation of prices for both industrial and domestic use.
At the same time, considerable governmental interest has
been shown in the development of hydropower and in the competition
of this form of energy with that generated through steam,
using solid, liquid, or gaseous fuels.
Although not generally realized and appreciated, this tendency
toward governmental regulation of fuel production and price
appears to be rapidly approaching a concerted effort to formulate
a directed and coordinated program (11) of governmental supervision
and control of all our natural resources, including fuel and
water power.
This governmental interest apparently involves control of
production, allocation of quantity, and source of supply. Its
purpose, presumably, is that of conserving our natural resources,
preserving the better grades of fuel and those with limited reserves,
and encouraging the use of available lower-grade fuels as
well as their more efficient production or recovery.
Anthracite producers’ efforts parallel those to establish minimum
prices for bituminous coal in that field.
C o n c l u s io n
The location of industry with reference to its fuel supply, the
available competing fuels, and their relative prices in the market,
improvements in combustion equipment, and the increasing use
of power, all combine to complicate the problem of purchase and
efficient utilization of fuel.
The purchase and use of fuel is in a state of evolution, or flux,
and the change has been accelerated to a marked degree by
governmental effort to regulate both supply and price.
Fuel selection and evaluation is by no means simple. The
consumer’s efforts may well be directed toward a better understanding
of his needs, a consideration of available information,
and a knowledge of the contemplated conservation measures (11)
and extensive procedure gradually being evolved for the control
of both price and supply of our natural resources. A proper interpretation
and appraisal of these several factors is of primary importance
in the satisfactory solution of his fuelpurchase problem.
Producers have come to realize that the sale of fuel is essentially
an engineering problem, and that fuel preparation, to meet
market requirements, is as important as its physical and chemical
characteristics. Many producers also realize that low rather than
high minimum prices are preferable, and that the use of low-cost
fuels will increase to the exclusion of those which may be priced
higher without possessing a correspondingly greater potential
heat value.
There has been suggested a yardstick, or measure, for evaluating
fuels within and between plants and for various methods of
conversion. A r6sum6 of the procedure for determining the
proposed prices for solid fuels is included, with reference to the
possible effect of governmental regulation. Reference has also
been made to the manner of proper fuel selection and the possibility
of cooperation between the combustion-equipment manufacturer
and the consumer, in order to make possible a wider
choice and use of sizes, grades, and kinds of fuel, to the end that
the consumer may more economically purchase and more advantageously
utilize available fuels.
The evaluation of fuels represents:
1 An integrated opinion of the producers of fuels which
enter a common consuming market.
STONE—PURCHASE AND USE OF FUEL 649
2 The relative values of available fuels, as determined by the
consumer on the basis of conversion in his own plant and equipment
and under conditions which govern his actual power
production.
The evaluation by both may vary over a considerable range.
Nevertheless, the integration of all valuations in each instance,
measured one against the other, becomes, as it were, a fair
average value of the proper differentials between the various
kinds, qualities, and sizes of fuels. The required realization,
modified as may be by the price for competing fuels, will determine
the final level of prices for a fuel in a common consuming
market.
B IB LIO G R A PH Y
1 U. S. N ational Bitum inous Coal Commission, W eekly Coal
R eport No. 1084, W ashington, D. C., April 23, 1938.
2 National Bituminous Coal Act of 1937, U. S. 75th Congress,
H .R. Bill 4985.
3 Bituminous Coal Tables 1936-1937 and 1937-1938. Published
by the Division of Research and Statistics. F. G. Tryon director
of the N ational Bituminous Coal Commission. Statistics subsequently
incorporated in part in Bureau of Mines M inerals Yearbook.
4 Exhibits P-813 (F. G. Tryon and J. W. M cBride), P-822
(C. J. P o tter), and other exhibits of the Bitum inous Coal Division,
D epartm ent of the Interior, Docket No. 15, Final Hearing, W ashington,
D. C.
5 M inerals Year Book, 1939 (and other years), U. S. Bureau of
Mines, W ashington, D . C.
6 Electric Power Statistics, Federal Power Commission, W ashington,
D. C., 1937. (Production and capacity for strictly private
use not included.) ;
7 Association of American Railroads, R eports for 1919-1920
and 1938. Also, Statistics of Railways, Class 1— 1919-1920 and
1938, B ureau of Railway Economics, W ashington, D . C.
8 “ M onthly and Annual Production of Electricity, E tc .”
(steam), also “Production of E lectricity for Public Use” (hydropower),
U. S. Geological Survey, to and including M ay, 1936, subsequently
Federal Power Commission, W ashington, D . C.
9 "Factors Recommended for Consideration in the Selection of
Coal” (second edition) sponsored by the American Society for Testing
M aterials and published by National Association of Purchasing
Agents.
10 Exhibit P-1099 (F. G. Tryon), Bitum inous Coal Division,
Docket No. 15, W ashington, D . C.
11 "E nergy Resources and N ational Policy,” National Resources
Com mittee, W ashington, D . C., January, 1939.
12 Connally Petroleum Act, originally passed U. S. Congress,
Senate Bill No. 1190, February 22, 1935; extended Senate fiill No.
790, June 4, 1937; and again Senate Bill No. 1302, June 29, 1939,
for three years.
13 Proposed Cole Petroleum Act, U. S. Congress, H .R . Bill No.
7372, introduced by Rep. Cole of M aryland, July 26, 1939, a bill to
provide for cooperation with states to prevent waste of petroleum ;
to create an office of Petroleum Conservator; to amend the Act of
February 22, 1935, as amended; also, U. S. Congress H .R . Resolution
No. 290, introduced by Rep. Cole of M aryland, August 21, 1939.
Discussion
H. K. D e a n .9 The last few years have brought about several
changes in fuel-burning equipment and in boiler and furnace
design. In so far as coal is concerned, progress in burning is
definitely toward pulverized coal, burned in either slag-tap
furnaces on the one hand or in dry-ash furnaces on the other.
There is less tendency to design the intermediate type, many of
which are still giving considerable trouble.
A great deal is yet to be learned regarding the characteristics
of ash, since its behavior is different for different methods of
burning. The standard ash-fusion-temperature determination
has, in the past, been given as of a reducing atmosphere and there
is a difference in the fusing temperature in an oxidizing atmosphere;
as much as 400 deg with certain types of coal.
9 Sales Engineer, The Babcock & Wilcox Com pany, Boston, Mass.
Mem. A.S.M .E.

650 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
A wide range of fuel utilization may be had, particularly in
burning fuels in suspension. A combination of pulverized coal,
oil, and gas can be and has been burned successfully in either the
slag-tap or dry-ash types of pulverized-coal furnace.
Changes in coal prices and marketing practices may alter the
natural movement of fuels as it is now known. Therefore, there
appears to be a greater necessity for full investigation and
accurate knowledge of the available fuel supply by the purchaser,
the engineer, and the manufacturer. Design characteristics
of steam-generating equipment and over-all economics
will probably place limitations on fuel purchases for some time
to come. There will probably always be certain characteristic
fuels which will give better over-all performance in certain types
of steam-generating equipment.
A. W. T h o r s o n . 10 In view of the proposed minimum prices
for bituminous coal, everyone concerned with the purchase or
utilization of this type of fuel is vitally interested in and should
make an exhaustive study of this paper. Very little engineering
information regarding the method of developing prices has been
available to date. The paper gives detailed information as to
the method of studying markets and production costs together
with the correlation of the two to set up the price schedules. If
and when these prices become effective, the value of this paper
will become more and more apparent.
Has the author prepared a chart similar to that given in Fig. 4,
but showing delivered prices at the selected destinations rather
than f.o.b. mine prices It would appear that such an exhibit
would be of added value to coal consumers.
A u t h o r ’s C l o s u r e
Mr. Dean and Mr. Thorson in their discussions have touched
upon two factors, transportation and conversion, which together
with the initial fuel cost at its source of supply have a fundamental
and important bearing upon available fuels, sources of
supply, and the markets which they can economically serve.
As I have indicated, the cost of transporting a ton of bituminous
coal by rail averaged, for the entire United States, an amount
equal to or greater than the cost of production. Coal, depending
upon the mine location and the market destination, may and often
does move in whole or in part by other competing forms of transportation,
that is, via lake, tidewater, river, and truck. Competition
between carriers and, what is perhaps even more impor­
10 The Chesapeake and Ohio Railway Com pany, D etroit, Mich.
Mem. A.S.M .E.
tant, the actual transportation rate itself may have an important
bearing upon the available sources of fuel supply. To a large
extent the cost of transportation may and does determine the
producer’s ability to retain or extend his fuel market, and it is not
unreasonable to expect that a reduction in transportation cost
might well result in a substantial increase in gross revenue to the
coal carrier.
The delivered cost per million Btu, other factors being equal,
becomes the consumer’s measure of purchased-fuel value, and of
this transportation forms a substantial part, in fact in many
consuming markets it amounts to several times the actual fuel
cost f.o.b. cars at the mine.
Indications are that in many instances the minimum prices
now proposed will be revised somewhat in the final price schedules,
and that consideration is being given to the modification
of transportation rates. For these reasons it would seem more
appropriate to defer for the present a graphic illustration of typical
delivered fuel prices, those in which the consumer is primarily
interested.
In the past, steam-generating equipment generally has been
geared to the available fuel supply, its characteristics, prices, and
marketing practices. Considerable progress has been made, as
Mr. Dean points out, in widening the range of fuel utilization and
particularly where fuels are burned in suspension, both singly and
in combination. While it is true that design characteristics of
steam-generating equipment and over-all economics may place
limitations on fuel purchases in many cases and for some time to
come, progress in the design of steam-generating equipment, the
technique of fuel burning, equipment maintenance, and ash or
waste disposal has been sufficiently great in the last few years to
indicate that steam-generating equipment having a much wider
range of fuel utilization can now be manufactured and installed
with very little additional effort and expense.
Thus, through improvements in transportation and utilization
of fuel, will the consumer be enabled to take greater advantage
of potential and changing fuel markets and be governed in his
purchases to a greater degree by the actual fuel value as measured
by the delivered cost per million Btu.
The producer, the carrier, the equipment manufacturer, and
the consumer, each have a real interest in the contribution of the
others in the fair evaluation, economical purchase, and advantageous
utilization of available fuels.
The expressed interest of Messrs. Dean and Thorson, as well
as that of many others, is encouraging and very much appreciated.

Salt-V elocity M easurem ents a t L ow
Velocities in Pipes
By LESLIE J. HOOPER,1 WORCESTER, MASS.
This paper describes tests made at the Alden Hydraulic
Laboratory to determ ine the performance of the saltvelocity
m ethod at low velocities. Three series of tests
were made: the first in a 40-in. pipe, the second in a 12-in.
vertical pipe in which the test section and its approaches
could be readily reversed, and the third in a horizontal 2-
in. pipe. The tests show that for flow in a long straight
pipe and for velocities lower than those norm ally found in
practice, there exists a critical mixing velocity below which
good mixing of the injected brine does not occur. An
analysis of the results indicates some of the factors which
affect the critical m ixing velocity. Finally, it was found
that the effect o f gravity on the accuracy of the m ethod,
when applied to a vertical pipe, is negligible as long as
proper mixing is secured.
THE salt-velocity method of water measurement has been
primarily used for the measurement of discharge in the
determination of pump and water-wheel efficiencies. In
these applications, penstock velocities are usually of the order
of one pipe diameter per second and no velocity limitations of
the method were indicated in this experience. The object of this
study was to investigate the sources of error at low velocities in
the application of the salt-velocity method to long straight pipes.
D e s c r i p t i o n o f A p p a r a t u s
This investigation was conducted at the Alden Hydraulic
1 Assistant Professor of Hydraulic Engineering, W orcester Polytechnic
Institute. Mem. A.S.M.E.
C ontributed by the H ydraulic Division and presented at the
Spring Meeting, W orcester, Mass., M ay 1-3, 1940, of T h e A m e r i c a n
S o c i e t y o f M e c h a n i c a l E n g i n e e r s .
N o t e : Statem ents and opinions advanced in papers are to be
understood as individual expressions of their authors, and not those
of the Society.
Laboratory of the Worcester Polytechnic Institute and the tests
were made during 1939 and early 1940.
A straight length of 40-in. pipe was selected for the test section,
Fig. 1. Four 3/ 4-in. pop injection valves were located at the introduction
station. The first set of electrodes was located 7.9 ft
downstream from the pop valves and the second set of electrodes
located 11.42 ft beyond the first set. The brine was injected under
30 to 50 lb per sq in. excess pressure, and the discharge of the
valves was controlled by a quick-acting lever valve in the brine
supply line. The discharge of the 40-in. pipe passed through a
36 X 16-in. venturi meter, thence through a 100-hp water
wheel and, finally, over an 8-ft suppressed weir in the tailrace.
The venturi meter and the weir were both calibrated just before
the tests by means of a 50,000-lb weighing tank.
The pop valves were built up from pipe fittings. Each
valve had a spring-loaded flat disk operating against a stop.
The valves required a pressure of about 5 lb per sq in. to open
them. The excess-brine pressure of 30 to 50 lb per sq in. caused
all the valves to be operated simultaneously and the springloaded
disks kept brine leakage at a minimum between shots.
Each electrode in the 40-in. pipe was a grid, consisting of nine
parallel Vs X 1/ 2-in. steel bars, carried on a wooden streamlined
center strut which was fixed across one diameter of the
penstock, Fig. 1. These steel bars were placed edgewise to the
flow to reduce their resistance and were arranged on 4 7/«-in.
centers across the pipe, with the first and last bars 2 7/s> in. from
the pipe wall. The bars were cut so as to leave an end clearance
of Vi in. The inside walls of all pipes tested were coated with
insulating paint. Alternate bars were connected together and
only one pair of leads was brought out from each electrode station.
In order to determine the effect of gravity on the heavier salt
solution when injected into a pipe line, the apparatus which is
illustrated in Fig. 2 was installed. The 12-in. vertical pipe with its
test section and approaches was so arranged that it could be turned
651

652 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
end for end, thus reversing the effect of gravity on the flow.
The four pop valves were located 18 in. from the flange of the
upstream elbow and the electrodes were located 3 ft 7 in. and
18 ft 7 in., respectively, from the pop valves. These injection
valves were built of 1/»-in. pipe fittings in the same manner as
the Vr-in. valves described. The electrodes consisted of two
The test section in the 2-in. line was located in a straight length
of standard pipe. The brine was injected through a flush
nipple located 4.61 ft downstream from a right-angle elbow.
The distance from the introduction station to the first electrode
was 2.39 ft and to the second electrode 10.48 ft. The electrodes
consisted of Vr-in. mesh galvanized-iron screen wire similar to
that used in the 12-in. pipe, except that the distance of separation
was Vs in. The discharge was measured in a 600-lb weighing
tank. A Bristol 5-amp strip-chart ammeter was used to record the
passage of the salt solution through the electrodes in the 40-in.
and 12-in. fine tests. A General Electric strip-chart recording polyphase
wattmeter was used in the 2-in. line tests. The power supply
for the electrode circuits was taken from a variable-voltage transformer.
The electrodes were connected in parallel to the
transformer with the recording instrument located in series in
the common lead. A time record was made on the recording chart
by a chronographic pen actuated from a calibrated timer.
The brine used in these tests was mixed to various specific
gravities as determined by means of hydrometers. For most of
the tests, solutions of common salt were used but dilute hydrochloric
acid was used for the 1.02 specific-gravity tests in the 12-
in. line. A hand force pump was used to inject the brine in the
12-in. and 2-in. pipe tests. The acid solution was injected under
air pressure with the pop-valve operation controlled by a quickacting
valve.
The volume between the planes of the electrodes was determined
by actual measurement of the diameter and length in
the 40-in. pipe. Four diameter readings were taken at every foot
along the test section together with 18 measurements of its
length. From these measurements the test volume was computed.
In the 2-in. and 12-in. lines, the test section was blanked off and
filled with water, which was then drained into a weighing tank.
This measurement was repeated 8 times and the test volume was
computed from the average weight and density of the water.
F ig . 2
T e s t S e c t i o n in 1 2 -In . P ip e
P r o c e d u r e
The procedure was essentially the same for all of the tests.
The discharge was adjusted to a desired value and from 5 to 10
min were allowed to elapse to secure steady flow conditions.
Then from 5 to 15 salt shots were recorded at the same time that
the discharge was being measured by the check method. For the
40-in. line the venturi deflection and the head on the weir were
read at intervals during each test. For the 2-in. and 12-in. lines,
the discharge was measured with the weighing tank.
C a l c u l a t io n s
The passage time for a salt shot was taken as the time in seconds
between the centers of gravity of the two electrode curves. The
volume of the pipe between the planes of the two electrodes divided
by the average passage time of all the salt shots in a run
gave the salt-velocity discharge.
The true discharge was that measured by the weighing tank
for the 2-in. and 12-in. pipe tests or the average of the venturi
and weir discharges for the 40-in. fine test.
The difference between the true discharge and that of the saltvelocity
measurement, expressed in per cent of the true discharge,
was termed the error of the salt-velocity measurement.
The mean velocity was found by dividing the true discharge
by the area of the pipe.
R e s u l t s o f T e s t s
The results of these tests are shown in tabular and plotted form.
In a long pipe, where the normal flow pattern has been established,
it is impracticable to introduce the salt solution uniformly
by means of pop valves or any other known arrangement. In

HOOPER—SALT-VELOCITY MEASUREMENTS AT LOW VELOCITIES IN PIPES 653
this case, the turbulence existing in tne pipe is reiiea upon to
secure uniform mixing of the salt solution with the flowing water
before the first electrode station is reached. With this in mind
it might be supposed that good mixing would be secured as long
as the flow in the pipe was in the turbulent range, as indicated
by a Reynolds number of 2500 or more. However, it has been
found by von K&rm&n and others that the turbulence existing
in a pipe may be materially reduced by the presence of an unmixed
solution either heavier or lighter than water.
To visualize this phenomenon, let us assume the worst case
where a lighter liquid is overlying a heavier liquid and they are
both traveling along a horizontal pipe together. For the present
■discussion, chemical or molecular diffusion may be neglected,
since these diffusion velocities are negligible when compared to
ordinary mixing velocities. If the two liquids are to be mixed,
part of the light liquid must move down into the heavier liquid
and a corresponding amount of the heavy liquid must be moved
upward. Then work is done in submerging particles of light
fluid in the heavier fluid and also in lifting the heavy particles into
the lighter fluid. Clearly, in the example selected, the axial
component of the fluid velocity has no effect upon the mixing

654 TRANSACTIONS OF TH E A.S.M.E. NOVEMBER, 1940
process and only the radial velocity components are effective in
promoting mixture of the two liquids. It naturally follows
that the work of mixing the two liquids is accomplished at
the expense of the energy available in the radial components
of flow. If the work required for mixing is less than the amount of
energy available, then the mixing process goes forward rapidly.
On the other hand, if the work required for mixing is greater than
the turbulent energy present, the process continues until the
energy available is used up and then, having reduced the crossflow,
the mixing process is checked. Thus, in this second case,
the presence of the heavier liquid tends to smother the crossflow
in the pipe.
In the example selected, the mixing was accomplished by the
radial components of flow but, in the more general case where
there is a small isolated mass of heavier liquid, mixing can take
place in all directions and it is more correct to say that the mixing
process depends upon the energy available in the fluctuating
components of the velocity.
From the foregoing, it would be expected that the change from
good to poor mixing conditions would take place rapidly. That
is, as soon as the heavier liquid starts mixing, its specific gravity is
reduced so that the process continues more easily. However,
if the heavier liquid is not diluted, it continues to smother the
fluctuating component of flow.
The work of mixing is accomplished at the expense of the
energy in the fluctuating component. The size of the average
fluctuating component is usually expressed as a percentage of the
average forward velocity and this is called percentage of turbulence.
For instance, 10 per cent turbulence means that the
average fluctuating component is '/io of the average forward
velocity at that point. Returning to the mixing process, when
a given shot of salt solution is injected into a pipe, a certain
amount of energy is required for mixing. Since the energy available
for mixing varies as the cube of the fluctuating component
and the fluctuating component is proportional to the forward
velocity in a given pipe, it follows that the energy available for
mixing varies as the cube of the forward velocity. Thus a small
change in the forward velocity results in a large change in the energy
available for mixing and it would be expected that the
change from good to poor mixing conditions would take place
suddenly. The test results show this to be true. The curves in
Fig. 4 show small experimental or casual errors at high velocities
which continue unchanged to some definite low velocity. Then
there is a “break” in the error curve showing a sudden departure
from the purely experimental type of error. The value of the
velocity at the break in the error curve, which is the velocity required
for good mixing, is termed the “critical mixing velocity.”
F u r t h e r T e s t s W i t h B r i n e a n d H y d r o c h l o r ic -A c id
S o l u t io n s
When brine is injected into water, the excess weight of the
brine is proportional to the difference of the specific gravities
of the brine and the water. Thus, brine of 1.10 specific gravity
has 10 times more weight in water than brine of 1.01 specific
gravity. The term “specific gravity minus one” is called the
effective specific gravity during the remainder of this discussion.
Tests were then made to determine experimentally the effect
of a change in the specific gravity of the injected brine upon the
critical mixing velocity. This was done first in the 12-in. pipe
and, subsequently, in the 2-in. pipe. In the 12-in. pipe test,

HOOPER—SALT-VELOCITY MEASUREMENTS AT LOW VELOCITIES IN PIPES 655
brine of 1.2 and 1.1 specific gravity and dilute hydrochloric acid
of 1.02 specific gravity were used. The results of the first two
tests came out so close together that a single curve represents
their average, Fig. 5. The test with the dilute acid, Fig. 6, indicated
that the solution was mixed at a lower velocity in the case
of upward flow but there was little change with the downward
flow. Furthermore, these test points show much more scatter
than in any of the other tests with brine. It was believed that the
acid probably produced secondary chemical and, therefore, electrical
effects at the electrodes which were zinc-covered, with the
consequence that the results were not as reliable as the tests made
with brine.
The results of tests with brines of 1.01 and 1.1 specific gravity
are shown in Fig. 4. Selecting an arbitrary error of —'/a per cent
in the salt-velocity measurement, a velocity of about 0.94 fps is
found with brine of 1.1 specific gravity and about 0.45 fps with
brine of 1.01 specific gravity. These results are in substantial
agreement with the mixing theory stated. On the one hand, the
work done varies directly with the “effective specific gravity” of
the solution being mixed, while on the other, the energy passing
through a unit area in the pipe cross section varies as the cube
F i o . 6
T e s t s i n 1 2 - I n . V e r t i c a l P i p e W i t h D i l u t e
H y d r o c h l o r i c A c i d
(A, downward flow; B, upw ard flow; specific g ravity 1.02.)
of the velocity. If the turbulent energy required is reduced to
Vio of its previous value, the pipe velocity required to produce
this energy is only reduced by (0.1)1/ ' or 0.465 of its previous
value. This is practically the change in the critical mixing
velocity found in the 2-in. line tests with a 10:1 change in the
“effective specific gravity” of the injected salt solution.
The total work done in mixing is also proportional to the
quantity of brine injected into a given volume of water flowing
in a pipe. The power required for commercial recording instruments
makes it difficult to secure sufficient electrode effect in
small pipes and thus the quantity of brine is usually increased
so that suitable curves will be obtained on the recording chart.
For instance, in the 2-in. line tests with the weak brine of 1.01
specific gravity, the ratio of the brine to the discharge was nearly
20 times greater than in larger pipes. Therefore, the critical mixing
velocities found in the 2-in. line tests are greater than those
necessary in larger pipes with the same flow condition.
The laboratory tests do not indicate directly the effect of wall
roughness upon the critical mixing velocity but an idea of its influence
may be obtained from theoretical considerations. The
previous discussion indicates that a definite value of the average
fluctuating component was necessary to produce satisfactory
mixing for a given flow condition. This value of the fluctuating
component may be obtained with a low percentage of turbulence
and a high forward velocity or with a high percentage of turbulence
and a correspondingly lower forward velocity. The present
theory of turbulence indicates that the percentage of turbulence
at a point in a pipe is proportional to the wall roughness. Therefore,
the greater the pipe roughness, the lower the pipe velocity
necessary to provide good mixing.
E f f e c t o f G r a v it y U p o n I n j e c t e d B r i n e S o l u t io n s
The effect of the force of gravity upon the injected brine or the
tendency of the brine to sink through vertically flowing water was
studied in the 12-in. pipe tests. At very low velocities when
proper mixing is not secured, the errors in the salt-velocity measurements
indicate that the unmixed brine masses tend to slip
through the flowing water under the influence of gravity. That is,
when the flow is upward, the salt masses travel more slowly
than the main body of water, resulting in a longer passage time
and an indicated discharge which is too small. With downward
flow, the brine tends to slip along faster than the water with a
resulting discharge which is too large. But a study of the curves,
Fig. 5, will show that, when there is proper mixing of the brine,
there is no systematic error of the salt-velocity measurements.
The importance of this fact cannot be overemphasized.
The test curves made with brine of 1.1 and 1.2 specific gravity
in the 12-in. vertical pipe are not symmetrical with respect to
the axis of zero error, the upward-flow tests showing a higher
critical mixing velocity than the downward-flow tests. It is
probable that, in the case of downward flow, the force of gravity,
acting upon the unmixed brine masses, tends to give them a
slightly increased velocity. This energy is dissipated in eddies
and, hence, promotes the mixing process. Conversely, with upward
flow, more energy must be supplied from the water flowing
in the pipe to lift the unmixed brine in addition to mixing it.
A p p l y i n g S a l t -V e l o c it y M e t h o d i n W a t e r - W h e e l I n t a k e
In this discussion of tests at low velocities with the saltvelocity
method, only the application of the method to a long
straight pipe has been considered. This assumes that normal
pipe flow exists at the point where the brine is injected and continues
through the measuring section. It might be inferred from
these tests that it would be impossible to apply the salt-velocity
method in the intake of a large water wheel where pipe flow
does not exist and, hence, where turbulence due to upstream wall
roughness is lacking. But there are two very important differences
to be taken into account When considering flow into an intake.
Since pipe flow has not yet started, all of the water is moving
forward with very nearly uniform velocity and, hence, the
nonuniform distribution of the brine, as it leaves the pop valves,
causes an insignificant timing error. On the other hand, turbulence
is probably considerably greater than would normally
be found at the entrance to a pipe due to the disturbing influence
of the racks which are always present in a water-wheel installation.
During 20 years of experience with this method, a large
number of intake-test installations have been made. In not one
of these instances was there any evidence of a lack of proper
mixing of the brine, although in some cases the water velocity at

656 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
the injection station was below that indicated in these tests as
being necessary for proper mixing. Thus, it is believed that the
factors concerning the critical mixing velocity which have been
discussed only apply to the measurement of flow in long straight
pipes with very low velocities.
S u m m a r y
The following conclusions may be drawn from these saltvelocity
measurements at low pipe velocities:
1 The low-velocity effects discussed in this paper are found
at velocities below those normally encountered in practice.
2 The critical mixing velocity is a very definite limiting
velocity below which proper mixing of the injected brine is not
secured.
3 The force of gravity does not have any influence upon the
accuracy of the salt-velocity measurement as long as the injected
brine is readily mixed with the water flowing in the pipe.
A c k n o w l e d g m e n t
The author gratefully acknowledges the assistance given by
Dr. Th. von K&rm&n who suggested the mechanics of the
mixing process and thereby contributed greatly to the value of
this investigation.
B IB LIO G R A PH Y
1 Basic description of the salt-velocity m ethod: “ The Salt-
Velocity M ethod of W ater M easurem ent,” by Charles M. Allen and
Edwin A. Taylor, Trans. A.S.M .E., vol. 45, 1923, pp. 285-341.
2 Study of the mixing process: “ How W ater Flows in a Pipe
Line,” by Charles M. Allen, Mechanical Engineering, vol. 56, Feb.,
1934, pp. 81-84.
3 Low-velocity study in open flumes: “ Contribution k l’Etude
de la Mesure des D ebits d ’E au par la M ethode Allen,” by M artin
Mason, Thesis at the U niversity of Grenoble, Revue GirUrale de
VHydraulique, Paris, 14.
4 Theoretical discussion of turbulence: “ Fluid Mechanics for
Hydraulic Engineers,” by H unter Rouse, McGraw-Hill Book Company,
Inc., New York, N. Y., 1938.
5 M easurem ents of turbulence in w ater: “ Investigation of
Errors of P ito t Tubes,” by C. W. H ubbard, Trans. A.S.M .E., vol. 61,
1939, pp. 477—492.
D iscussion
D. P. B a r n e s .2 Referring to this and the paper3 by O. H.
Dodkin, it is begining to be recognized that, for the salt-velocity
method to yield ideally correct results, it would be necessary
that all the small elementary volumes of salt solution spend
statistically equal periods of time in each of the different velocity
zones from the center line to the outer radius in passing between
electrode stations. This is to say that turbulent mixing should
be continuous and complete transversely across the pipe throughout
the passage. Any deviation from this continuous symmetry
and uniformity of mixing would mean that part of the salt cloud
must lag behind the mean flow, the center of gravity of the salt
cloud, therefore, with it.
The natural tendency of a dense solution to settle downward
and to underflow a lighter liquid as an integral body is now a well-
established experimental fact. In the salt-velocity test, if this
tendency is not completely countered by the turbulent mixing, it
would be expected that in time the salt cloud would become progressively
less symmetrical or less uniform with respect to the
pipe.
In a sloping pipe, the effect of the tendency for the denser
2 California In stitu te of Technology, Pasadena, Calif.
3 This is a joint discussion of the L. J. Hooper paper, and the paper
“ Field Checks of the Salt-Velocity M ethod,” by O. H . Dodkin, published
on page 663 of this issue of the Transactions.
liquid to assume an asymmetrical distribution is compounded
with the effect of the component of underflow in the downhill direction.
In any case the asymmetry or nonuniformity alone
would tend to produce a velocity reading smaller than the true
mean, and this effect would be increased in the case of pumps with
uphill flow or decreased in the case of turbines with downhill flow
by the downhill component of the underflow.
The numerous field checks would appear to establish that the
errors introduced by these tendencies are slight and, that where
the method is used with full recognition of the possible sources of
error and where a proper allowance is made for them, the results
may be supposed accurate within perhaps 0.5 per cent, at least
for downhill flow. This is a view to be accepted with caution,
however, for bends, obstructions, steep slopes, low velocities, or
abnormal velocity distributions in the penstocks must continue
to introduce uncertainties wherever their effects are not subject
to calibration. It would seem that each proposed application of
the method should therefore be considered largely on its own
merits.
A n d r e w F e j e r 4 a n d J. W. D a il y .5 During the last few years,
it has become more and more desirable to determine large rates of
flow with great accuracy. In the case of large-scale field tests the
salt-velocity method has been found particularly suited for such
measurements because of its simplicity and convenience in application.
Very little is known, however, about the performance of
the method at low velocities, and this paper furnishes useful experimental
information in that range.
In connection with investigations conducted at the hydraulicmachinery
laboratory of the California Institute of Technology
on the design, operation, and testing of large-capacity pumps and
especially in connection with the question of extrapolation of
model results to prototype results, it has been of considerable interest
for the Institute to subject the salt-velocity method to experimental
study. The experimentation was assigned to the
writers, who present the following summary of results with the
approval of the laboratory directors.
E x p e r im e n t a l W o r k
Following Dr. von ICdrmdn’s suggestion that a salt cloud would
tend to settle downward in water due to its higher density whenever
the turbulent forces were inadequate to keep the cloud in
suspension, it was decided to obtain qualitative information about
shape and character of salt clouds under different flow conditions.
The tests undertaken were made in a straight, transparent,
lucite pipe 20 ft long and 5'A in. inside diam, mounted in such a
way that it was possible to change its inclination from 0 to 13 deg.
Solutions of sodium nitrate deeply colored with potassium permanganate
were injected into the flow through pop valves under
compressed-air pressure, and visual, photographic and cinematographic
observations of the behavior of the cloud on its path
through the pipe were made. Two types of injection nozzles
were used, i.e., a single central pop valve in the pipe axis and a
combination of six pop valves distributed uniformly over the
cross section. Solutions of five different densities (1.2,1.15,1.10,
1.05, 1.025) were injected into flows of velocities, covering a range
of 0.1 to 1.5 fps. This range corresponds to a range of 0.5 to 7.5
fps for geometrically similar flow in a pipe 10 ft in diam.
D is c u s s io n o f R e s u l t s
It was found that the salt cloud retained its symmetrical shape
only at considerable flow velocities. This can happen only if the
4 Research Assistant, Hydraulic M achinery Laboratory, California
In stitu te of Technology, Pasadena, Calif.
6 Research Fellow, H ydraulic M achinery Laboratory, California
In stitu te of Technology, Pasadena, Calif. Jun. A.S.M.E.

HOOPER—SALT-VELOCITY MEASUREMENTS AT LOW VELOCITIES IN PIPES 657
(b)
(a)
F ig . 7 S a l t C l o u d ; S p e c if ic G r a v it y o f B r in e 1.025; M e a n V e l o c it y o f F lo w 0.1 F p s ; F r o u d e N u m b e r 0.03
(6) (a)
F i g . 8 S a l t C l o u d ; S p e c if ic G ra v it y o f B r in e 1.025; M e a n V e l o c it y o f F l o w 0.5 F p s ; F r o u d e N u m b e r 0.6
(M (a)
F ig . 9 S a l t C l o u d ; S p e c if ic G r a v i t y o f B r i n e 1.15; M e a n V e l o c i t y o f F l o w 0.3 F p s ; F r o u d e N u m b e r 0.04
(6) (a)
F ig . 10 S a l t C l o u d ; S p e c if ic G r a v it y o f B r in e 1.15; M e a n V e l o c it y o f F lo w 1.1 F p s ; F r o u d e N u m b e r 0.55
energy of turbulent fluctuation is sufficient to overcome the differential-gravity
forces. It may be considered that for a given
setup the scale of the turbulence is independent of the rate of
flow, while the magnitude of the turbulent-velocity fluctuations
is proportional to the mean velocity of flow. This means that
the energy of the turbulent fluctuations is proportional to the
square of the velocity of flow. For this assumption, the ratio
between the energy of turbulence and the work done by differential
gravity forces may be expressed by a dimensionless quantity
which contains in the numerator the square of the mean velocity
and in the denominator the product of the effective gravity and
a linear dimension of the setup (say the diameter of the pipe).
This quantity, which has the form of a Froude number for a particular
setup, is written
ever the density of the injected solution is increased or the velocity
of flow decreased so that the Froude number is smaller
than the critical, a secondary or density-flow current is set up,
the cloud sinks to the bottom of the pipe, and the assumption
of the salt-velocity method that the average velocity of flow equals
the average velocity of the cloud is no longer valid. Figs. 7 to 10
of this discussion8 show typical salt clouds within a range of
Froude numbers from 0.03 to 0.68. Figs. 8 and 10 show clouds
corresponding to the critical Froude number which was found to
be reasonably constant over the range of velocities and densities
considered. For tests with the single nozzle / ’’em = 0.7, and
with the multiple nozzle F„it = 0.56. The difference is due to
the range in the scale of turbulence caused by the change in the
geometry of the flow.
R e m a r k s o n t h e F r o u d e N u m b e r
The Froude number just discussed should express the ratio of
the inertia forces to the effective gravity forces acting on the
cloud shortly after the effects of the injection have disappeared.
As stated, the term p2 should be the actual “initial cloud density.”
For practical use of the Froude-number criterion, it is necessary
6 Figs. 7(a) to 10(a) show the clouds immediately after the injection,
while Figs. 7(b) to 10(6) illustrate the same clouds some distance
downstream. The numbers give the traveled distance from the injection
nozzle in pipe diameters.

658 TRANSACTIONS OF TH E A.S.M.E. NOVEMBER, 1940
to correlate this initial density with the known density of the injected
brine. Two simple assumptions can be considered as
limiting cases:
1 The brine suffers no volume change between the instant of
injection and the formation of the cloud; and
2 The injected brine is distributed immediately with a rate of
dilution proportional to the velocity of flow.
As Figs. 7(a) to 10(a) indicate, the volume of the brine cloud does
increase immediately after injection. On the other hand, the
second assumption is not strictly valid. Rough measurements
from photographic records show a definite tendency of the cloud
size to decrease with increasing velocity of the main flow. This
apparently is due to the fact that the penetration of the brine in
the lateral direction decreases with increasing velocity. For
high velocities, the cloud has an elongated shape with a small diameter
while for small velocities the cloud has a large diameter
even though of short axial length. It appears to the observer
that because of these facts the two effects approximately balance
each other. Consequently, the assumption of a constant “initial
cloud size” seems reasonably justified. Since the density of the
brine solution as injected is proportional to the “initial cloud
density” for a constant injected volume and a constant initial
cloud size, it was introduced into the expression for the Froude
number.
It should be noted that the critical Froude number is not necessarily
the same for different setups. The intensity and scale
of turbulence can be different for various setups even in the case
of geometric similarity. The type of the injection nozzles, the
size of the injected volume, and the pressure used for the injection
affect the initial cloud size. However, the experimental results
mentioned show that, for a fairly wide range of injection conditions,
the values of Fcrit vary between narrow limits.
C o n c l u s io n s
Roth the author’s work and the experiments noted in this discussion
appear to indicate conclusively that salt clouds, as used in
the salt-velocity method for rate-of-flow measurements, sink by
virtue of the differential gravity forces into regions of lower velocities
whenever the rate of flow and therewith the turbulent
forces decrease beyond a certain minimum value. Below this
velocity, which appears to be mainly dependent upon the Froude
number of the test, it cannot be assumed that the cloud indicates
the average velocity of flow.
It would be interesting to combine the author’s method with
the type of attack herein mentioned to determine whether the
breakdown of the method, as shown by his experiments, coincides
with the “critical Froude number” or whether additional sources
of error counteract or accentuate the effect of the density currents.
R. T. K n a p p .7 These first published results on one of the auxiliary
phenomena accompanying the salt-velocity method of water
measurement are of value, especially, since it is possible that the
effects of the density difference between the salt cloud and the
water may form one of the basic limitations on the accuracy of
the method.
The writer considers it rather unfortunate that the author restricted
his measurements to horizontal and vertical pipes. The
vertical pipe is not a typical case, since it represents a condition
of equilibrium between the.two fluids. I t is true that this equilibrium
is unstable, but it is probably very effective in the short time
interval which exists in a salt-velocity measurement. Indeed, the
existence of this equilibrium has been utilized by some Swedish
7 Associate Professor of H ydraulic Engineering, California Institu
te of Technology, Pasadena, Calif. Mem. A.S.M .E.
investigators8 in the development of a sedimentation type of
particle-size analyzer in which a high-density suspension is superimposed
upon a body of lower-density liquid. Thus, it seems
that it would be unsafe to consider a vertical pipe as an extreme
case of an inclined one and to use the results from the vertical
flows to predict what would happen in the inclined pipe.
It should be remembered that the effect to be anticipated, as
the result of the difference in density of the two flows, is basically
a lateral one and not an axial one. In the horizontal or inclined
pipe, the salt cloud tends to settle toward the lower side of the
pipe, thus effectively lowering the average velocity of the solution.
Of course, in the inclined pipe, there will be an axial component
as well, which will tend to increase or decrease the velocity,
depending upon the direction of flow. However, this
axial component is probably of secondary importance.
In conclusion, the writer feels that this phenomenon is a very
important one and hopes that the present paper represents only
the beginning of the author’s investigations on the subject.
M . A. M a s o n .9 The suggestion that the possibilities and applicability
of the Allen method might be defined by criteria relating
the physical and hydraulic characteristics of a proposed
test section10 has been applied to pipe flow by the author, and it
seems that at last the Allen method may be thoroughly studied
in a logical manner, looking to the complete definition of its place
in the art of water measurement.
The writer would, however, disagree with the author’s choice of
a criterion for good mixing. It is not believed that the ‘‘critical
mixing velocity” is a sufficient definition of the hydraulic characteristics
of a waterway to permit its use as a criterion of the mixing
conditions in the test section. A moment’s thought will show
that, for a particular system in which normal pipe flow has been
established, a critical velocity can be found which defines the
mixing ability of the flow under the conditions existing. However,
as stated by the author, “the percentage of turbulence at a
point in a pipe is proportional to the wall roughness. Therefore,
the greater the pipe roughness the lower the pipe velocity necessary
to provide good mixing.” It follows that the mean velocity
of flow is not a satisfactory criterion of mixing, except for specified
conditions which, in general, will apply only to the system
used to evaluate the criterion.
A more important objection to the use of a velocity parameter
as a eriterion lies in the omission of a time element defining the
period during which mixing is accomplished. To be of value in
the application of the method, the mixing criterion must define,
for a standard injection procedure, the minimum distance along
the waterway required for the complete mixing of injected solution
and water. Obviously the proposed velocity criterion does
not furnish sueh information and, in fact, defines only the mixing
ability of a flow under certain unique conditions.
From consideration of the mechanism of mixing and the character
of turbulent flow, one may specify the probable components
of a general criterion of mixing. The percentage of turbulence,
i.e., the proportion of the forward velocity of flow present as
fluctuating turbulence velocities in the flow, defines the ability
of the stream to transport solution from the point of injection to
other points in a cross section. If the turbulence is isotropic,
then the percentage of turbulence is therefore a measure of the
8 “ M easurem ent and Significance of G rain Size D istribution of
Cement Particles,” by D. W erner and S. Hedstrom , Zement, vol. 17,
p art 1, June 28, 1928, pp. 1002-1005; p art 2, July 12, 1928, pp.
1071-1076.
9 Engineer, Beach Erosion Board, W ar D epartm ent, Washington,
D. C. Jun. A.S.M .E.
10 “ Contribution h l’fetude de la mesure des debits d ’eau par la
methode Allen,” by M. A. Mason, Revue gSnerale de VHydraulique,
no. 28, July-Aug., 1939.

HOOPER—SALT-VELOCITY MEASUREMENTS AT LOW VELOCITIES IN PIPES 659
rate of diffusion of the solution throughout a cross section due to
turbulent mixing. The percentage of turbulence is certainly
one of the components of the criterion, and includes not only the
velocity, but also the roughness characteristics of the waterway.
Obviously, a size factor of the conduit must also be included
injthe criterion; for a closed channel the area, and for an open
channel, either the wetted area or the hydraulic radius might be
considered.
A third factor, time, must be included. Considering percentage
turbulence as a measure of the spatial rate of transport of
solution throughout the water, it is evident that a certain time
must elapse for the transport of a salt particle from, say, the center
of the conduit to the wall, or vice versa. The time factor
might then be expressed as the ratio of water depth or pipe diameter
to the radial component of the turbulence fluctuations.
The fractional portion of the linear dimension to be employed in
this expression will, of course, depend upon the injection pattern.
If a standard pattern of injection can be adopted the influence
of injection procedure may be eliminated, and more important,
the time factor may be very considerably reduced by the use of
a well-distributed injection pattern.
Assuming the foregoing analysis to be applicable and letting ut
be the mean turbulence fluctuation; A the cross-sectional area
of the channel; and T the time factor; the form of the mixing
criterion might be
M = u, A T ...................................... [1]
having the dimensions L 3 of a volume.
Because of the practical difficulty of measuring u„ it may be
desirable to substitute for u, the expression11 lv, where I is the
Prandtl mixing length, and v is the kinematic viscosity of the
water. Experiment alone will determine the most satisfactory
term to be used as a measure of the turbulent diffusion, the suggestions
given herein being intended principally to indicate the
nature of the term.
The results of the author’s tests to determine the effect of
gravity on the operation of the salt-velocity method are a further
corroboration of the writer’s thesis1 that turbulence in the stream
is the chief factor governing the accuracy of the method.
E. A. T a y l o r .12 The author has described a study which
gives a laboratory answer to a salt-velocity question which has
arisen several times in the field. This question was: “When salt
is introduced into a penstock does the higher specific gravity of
the salt solution drive the salt through the water and toward the
bottom of the penstock, thus changing the time of passage of the
salt solution through the test section, and introducing an error
in salt-velocity measurements”
At many of the high-head power plants, where salt-velocity
tests have been made, the penstocks have very steep slopes and,
at some of them, the penstocks approach the vertical for considerable
distances. If this difference in the specific-gravity question
were important, then the computations for turbine efficiency
at those plants might be appreciably in error.
The results of the laboratory studies described by the author
indicate that the effect of gravity on the salt-velocity method of
water measurement is negligible, when proper salt mixing and
distribution are secured.
In 1939, the salt-velocity method of water measurement was
used in making efficiency tests on the Colorado River Aqueduct
pumps for the Metropolitan Water District of Southern California.
During those tests, a field answer was found to the
specific-gravity question. The question was raised at the Iron
11 “ Modern Developments in Fluid Dynam ics,” edited by S.
Goldstein, Oxford University Press, New York, N. Y., 1938, chapter 5.
I! Consulting Engineer, Worcester, Mass.
Mountain Pumping Station and the district engineers agreed to
turn that plant into a field laboratory to study the subject.
The regular salt-velocity test section, used for pump-efficiency
tests at this plant, was 200 ft long, 100 ft from salt injection to
the first set of electrodes, and another 100 ft to the second set of
electrodes. For this specific-gravity study, two additional pairs
of electrodes were installed in the 10-ft delivery pipe.
These additional electrodes were called spot electrodes and
consisted of small steel plates, 4 in. square and spaced 2 in. apart.
These spot electrodes were fastened 5 ft beyond the regular second
set of electrodes, making the length of test section used for the
spot electrodes 105 instead of 100 ft. One pair of spot electrodes
was located at the top of the pipe and the other pair at the bottom.
During the study, the ammeter connections were alternated,
shot by shot, between the two sets of spot electrodes.
A total of 74 salt shots were made and computed in this study.
Averaging the results of these shots showed that the passage time
to the top spot electrodes was 0.04 sec less than the passage time
to the bottom electrodes. This was regarded as a perfect check
and a conclusive field answer to the gravity question.
As further assurance of the accuracy of the salt-velocity method
of water measurement and confirming the belief that any specificgravity
error was too small to be seriously considered, two volumetric
check tests have been made at Pacific Coast plants, one of
them being a high-head plant with a steep slope in the penstock.
For those check tests, a 3-acre forebay, 20 ft deep, and a 10-
acre reservoir, 10 ft deep, were used as basins for the volumetric
measurements. The volumetric results checked salt velocity,
by 0.75 per cent at one plant and by 0.5 per cent at the other
plant. The latter was the high-head steep-penstock plant where
any gravity effect would be expected to be most apparent.
The author’s conclusion, from laboratory tests, th at “for velocities
lower than those normally found in practice, there exists a
critical mixing velocity below which good mixing of the injected
brine does not occur” is confirmed by some field tests recently
made.
Field efficiency tests were made on the Colorado River Aqueduct
pumps in May and June, 1939. At Intake, the first station
tested, the discharge values for single pumps checked the discharges
when two or three pumps were operating together.
At the second station tested, Gene, the results of the preliminary
tests, comparing single- and multiple-pump discharges, did
not check.
At all five Colorado River Aqueduct pumping stations, one 10-ft
delivery pipe carries the discharge from the three pumps now installed
at each station. When only one pump is operating, the
velocity in the delivery pipe is comparatively slow, about 2 fps.
The inside of these delivery pipes is coated with an enamel finish,
as smooth as glazed porcelain, and there are no rivet heads or
other projections into the inside of the pipe.
This combination of extremely smooth pipe and low velocities
created a condition of low turbulence, too low for satisfactory
salt mixing, during the passage of a shot. This accounted for
the failure of the discharge values to check between single and
multiple pumps.
To provide more turbulence, an artificial agitator, called a “tur-
bulator” was installed beyond the salt-injection station. This
so-called turbulator was a steel disk, 6 ft in diam, fastened in the
center of the delivery pipe and normal to the axis of the pipe.
This disk had an 18 X 24-in. hole in the center and was located
40 ft beyond the pop valves. The turbulator was used
during the remainder of the pump tests and single-pump-discharge
results checked the multiple-pump values.
No artificial turbulator was required at Intake, because the
salt-injection station was just above the manifold where the three

660 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
6-ft pipes enter the one 10-ft delivery pipe. The pop valves were
only 25 ft beyond this manifold and the turbulence caused by the
bends in the pipes was sufficient to last throughout the saltvelocity
test section.
R. M. W a t s o n . 13 From information now available, it appears
that, with normal pipe roughness, mean water velocities, reasonable
test sections, and when handled by a competent test crew,
the salt-velocity method for determining the rate of flow of water
is consistently accurate within 0.5 per cent. The writer belives
this to be the limit, at present with few exceptions, to which any
evaluation should be made. This accuracy is quite comparable to
that by which the other factors, pressure and power, which express
the performance of hydraulic machinery, can be determined
reliably in most commercial installations.
Among the many factors which affect the accuracy of the saltvelocity
method are the velocity of flow and the degree of turbulence
in the stream. As shown by the author, a certain minimum
turbulence is necessary in order that the salt cloud will travel
with the flow without slippage. Apparently this problem had
not been encountered in the field until a recent installation which
had an exceptional combination of low velocities and smooth pipe
walls. Unexplainably low results indicated trouble which was
traced to the absence of turbulence. Excessive deviation from
the expected results was corrected by the introduction of a “turbulator,”
the purpose of which was to create high turbulence
ahead of the salt-injection nozzles.
In view of the difficulties experienced in the initial tests, opportunity
was taken tb make a check of the effect of the low velocities
at those stations at which pumps manufactured by the writer’s
company, were installed. Each station consisted of three pumps,
each pump discharging from a 6-ft pipe into a common 10-ft pipe.
The salt-velocity test section was in this 10-ft line, with approximately
4 pipe diameters between the turbulator and the saltinjection
nozzles; 10 pipe diameters between the injection valves
and the first electrodes, and 10 diameters between the first and
second electrodes. The 10-ft discharge pipe at the test section
ran upward at an angle of about 25 deg with the horizontal.
The tests were made by the customer’s crew of engineers.
This test crew had had the experience of previous salt-velocity
work, in addition to experience with the use of this method on
nine other pumps in three stations at this same large development.
In every case the same problem existed; that of measuring by the
salt-velocity method the rate of flow with very low mean pipe
velocities and with extremely smooth pipe walls. This same
crew was present and assisted in the development of a satisfactory
technique for the method under the same conditions of low
velocities and smooth pipes prevailing at the first three stations,
and was therefore well qualified to conduct similar tests under
similar discharge-pipe conditions at the two stations in which the
writer’s pumps were installed.
Piezometer taps were placed at two sections of different areas
in the suction cones of all three pumps. The reduction in areas
was sufficient to give a differential reading of approximately 6 in.
of mercury in the range of flows considered. These venturi
meters were calibrated individually by salt-velocity tests made
during the operation of each pump singly. The salt-velocity
test section was in the 10-ft pipe common to all three pumps.
Each pump was operated individually, then in pairs, and finally
all three together. Salt-velocity determinations of the flow were
made each time. These calibrations were used during the multipump
runs to measure the flow delivered by the individual pumps.
During the multipump run, the sum of the individual flows thus
13 Engineer, Centrifugal Pum p Division, W orthington Pum p and
M achinery Corporation, Harrison, N. J.
determined was compared with the total as measured by the saltvelocity
method. The individual pump flows varied from approximately
180 cfs to 215 cfs, with corresponding mean pipe velocities
varying between 2.3 fps and 2.7 fps. For two and three
pumps operating together, the mean velocities in the test section
were, respectively, 2 and 3 times the mean velocities in the test
section during individual pump operation.
As a result of an analysis of the data by the writer, it appeared
that the salt-velocity results obtained with the low mean pipe
velocities, existing during the individual pump runs, were low
by between 0.3 per cent and 0.6 per cent, as compared to the
salt-velocity results obtained with the higher mean pipe velocities
existing during the multipump runs. It is interesting to note
that the laboratory results presented by the author confirmed the
possibility, direction, and order of magnitude of this discrepancy.
It is equally interesting to note that the discrepancy between the
single-pump and multipump tests was within the 0.5 per cent
accuracy obtainable with the salt-velocity method under normally
favorable conditions. It was only the peculiar combination of
circumstances present at this installation which warranted the
attem pt for even better accuracy.
In general, the difficulties experienced in the measurement of the
flow by the salt-velocity method were traceable, as brought out
in the paper, to insufficient turbulence to prevent an actual slippage
and retardation of the salt cloud due to its greater density,
particularly immediately after the injection. This resulted in
an apparent rate of flow which was considerably less than the
actual. Improvement was obtained by installing a turbulencecreating
construction a short distance ahead of the injection
valves. The trend, however, of any error from this effect was
shown definitely in these field tests to be negative, as is indicated
by the laboratory tests of the author.
Even with the artificially created turbulence, the analysis of
the field tests on six of the pumps as made by the writer, indicated
that some slippage did occur. In explanation, it is probable that
the turbulence resulting from the constriction decreased sufficiently
within the test section to permit some slippage of the
slightly more dense salt cloud. Certainly such is possible if the
rather high turbulent energy in the flow leaving the pump (although
this energy would be in small-scale eddies), and the turbulence
created at the junction of the 6- and 10-ft pipes during singlepump
operation, decreases sufficiently before reaching the injection
nozzles to cause the troubles noted.
In opposition to the results of the writer’s analysis of field
tests and checks and resulting conclusions, conditions at one of
the other three stations of this project permitted a volumetric
check of the salt-velocity results. By this check the salt-velocity
results appeared approximately 0.5 per cent high, after
installation of the turbulence-creating construction. However,
the writer does not know the full details of this check, and cannot,
therefore, attem pt to analyze possible differences.
In connection with the field checks of the salt-velocity method,
the writer wonders whether or not an examination has disclosed
any common characteristics of settings, or other factors, which
tend to give high results in some cases, and low results in others.
Such an examination should help to predict the direction of any
error and possibly ultimately to eliminate some of the causes.
With the increasing size of centrifugal pumps and an expanding
knowledge of design, pump efficiencies have now reached and even
exceeded 90 per cent. The pump designer is therefore no longer
satisfied with the 1 per cent to 2 per cent over-all test accuracy
which used to be acceptable. The ability to differentiate between
the effects of small design changes requires increasing accuracy
in the tests as the efficiency increases. Whereas, at 80 per cent
pump efficiency an accuracy of 2 per cent is sufficient to detect

HOOPER—SALT-VELOCITY MEASUREMENTS AT LOW VELOCITIES IN PIPES 661
really significant design differences, this accuracy should be
better than 0.5 per cent when the efficiency approaches or exceeds
90 per cent.
A u t h o r ’s C l o s u r e
The discussion presented by Messrs. Fejer and Daily is of great
interest to the author, especially the suggestion in connection
with the Froude number. This number, however, cannot be used
as a guide until more tests have been made to determine its critical
limits and the factors affecting those limits, because it does not of
itself completely define the flow conditions. Furthermore, in
addition to the pipe velocity, pipe diameter, and brine specific
gravity, which factors make up the Froude number, the roughness
of the pipe wall has a direct bearing on the turbulence, as does also
the viscosity of the fluid and the quantity of brine injected per
second-foot of flow. Finally, there is a modifying factor introduced
by the performance of the brine after being injected.
The first three points were mentioned in the paper, but the last
may need further clarification. It has been shown by experiment
that the same amount of salt solution, injected at a constant
rate, provides practically a constant height of electrode
curve over a wide range of velocities. Since the duration of the
shot remains constant, it follows that for low velocities the salt
is confined to a small volume of the penstock water and is distributed
over a proportionally larger volume of penstock water at
higher velocities. Thus the total energy for mixing becomes
the product of the energy per pound and the quantity of water
flowing. This indicates experimentally that the energy available
for mixing varies approximately as the cube of the velocity, rather
than the square.
With these thoughts in mind, the critical Froude numbers are
presented for those tests where complete information is available,
as follows:
Pipe C ritical Brine,
T est diam , velocity, specific
no. ft fps g ravity F actor
1 0 .1 7 0 .4 5 1.10 3.70
2 0 .1 7 0.94 1.10 1.61
3 1.00 0 .9 0 1.10 0.25
4 3.33 0 .6 7 1.06 0.07
5 10.00 2 .5 0 1.18 0.11
C altech 0.42 . . . . 0 .5 6 -0 .0 7
The results with the 2-in. pipe seem to be far out of line but it
is known in test No. 1 that the quantity of brine introduced per
second-foot of flow was nearly 20 times larger than usual. Dividing
3.7 by 20 gives 0.19, which brings the result more in line with
the figures obtained for the larger pipes.
The values of critical Froude number obtained by actual saltvelocity
measurements are about */6 to '/s the values given by
Fejer and Daily and they extend over a rather wide band. At the
start of these tests, all of the factors controlling the mixing process
were not appreciated, so that some of the experimental information
is approximate. This applies particularly to the amount of
salt injected per second-foot of flow. It is hoped that some of
these tests can be repeated at an early date so that an experimental
verification of this Froude number may be obtained.
Referring to Mr. Mason’s discussion, it was not the intention
of the author to suggest the use of the “critical mixing velocity”
itself as a criterion for good mixing. As stated in the paper, the
mixing process is affected by (1) the amount of brine injected;
(2) the specific gravity of the brine; (3) the roughness of the pipe
wall; (4) the size of the pipe; and (5) the velocity of flow in the
pipe. All of these terms affect the mixing process in the first
power except the velocity factor which varies as the cube. Therefore,
a change in the velocity is more significant than a change in
any other factor. From this point of view, the term “critical
mixing velocity” was used in a descriptive way and not as a sole
criterion of performance.
Furthermore, Mr. Mason states that the time or mixing distance
should appear in the mixing criterion. The mixing distance
is a factor intimately connected with the design of the pop valves
and their installation with respect to the first electrodes; from
that point it is most important. But, it was pointed out in the
paper that, if the mixing process starts, it will continue; however,
if the injected brine smothers the turbulence, the brine falls
to the lower part of the pipe and remains segregated. Under
these conditions, doubling the mixing distance or time will not
improve the mixing. I t is believed, therefore, that time or distance
should not be represented in the mixing criterion.
With regard to Mr. Watson’s discussion, it may be true there
actually existed a small error for the flows measured for the singlepump
units. Three similar pumping plants were tested in the
same fashion by engineers of Professor Allen’s office. In each
case, the average agreement of the sum of the single pumps, compared
with two and three pumps operating together, was to 0.1
per cent. This does not mean that the over-all accuracy of the
method is 0.1 per cent on a field test.
In conclusion, it is realized that the work reported in the paper
has only been a start in the investigation of the salt-mixing process
at low pipe velocities and the constructive discussions which
have been offered are appreciated.

Fio. 1
S a l t - V e l o c i t y L a y o u t a t U p p e r E n d o f S e r r a P e n s t o c k
Field Checks of the Salt-V elocity M ethod
By OSWALD H. DODKIN,1 SAO PAULO, BRAZIL
This paper gives the results of som e tests made in Brazil
on which verification of the reliability of the salt-velocity
m ethod of measuring discharge was obtained by cross
checks w ith the m ethod itself, by check tests with the
Gibson pressure-time m ethod, and by volum etric tests.
These show a highly satisfactory agreem ent even in setups
where the testing layout was far from ideal.
THE problem of measuring accurately the large discharges
of modern hydraulic turbines is one that has engaged
the efforts of many hydraulic engineers. About two
decades ago the Gibson pressure-time method2 and the Allen
salt-velocity method5 were introduced to the profession. Both
possessed advantages of simplicity and ease of application as
compared with some of the older methods. Laboratory tests
1 H ydraulic Engineer, The SSo Paulo Tram way Light and Power
Company, Ltd.
2 “ The Gibson M ethod and Apparatus for Measuring the Flow
of W ater in Closed Conduits,” by N. R. Gibson, Trans. A.S.M .E.,
vol. 45, 1923, pp. 343-376.
3 “ The Salt-Velocity M ethod of W ater M easurem ent,” by C. M.
Allen and E. A. Taylor, Trans. A.S.M.E., vol. 45, 1923, pp. 285-341.
C ontributed by the Hydraulic Division and presented at the
Spring Meeting, Worcester, Mass., M ay 1-3, 1940, of T h e A m e r i c a n
S o c i e t y o f M e c h a n i c a l E n g i n e e r s .
N o t e : Statem ents and opinions advanced in papers are to be
understood as individual expressions of their authors, and not those
of the Society.
and some field tests showed also a very promising degree of accuracy.
Since then considerable experience has been obtained
with these methods which in general has confirmed the early
promises of accuracy of both methods.
Both the Gibson and salt-velocity methods have been used with
success in the several plants of the companies subsidiary to the
Brazilian Traction, Light and Power Company, Ltd., the former
since 1926 and the latter since 1924. The purpose of this paper
is to report verifications of both methods obtained in these tests.
There will be given two tests in which the accuracy of the saltvelocity
method is confirmed by itself with a different apparatus
setup, two tests in which the results are confirmed by the Gibson
method, and one in which its accuracy is confirmed by volumetric
test.
S e r r a U n i t 5
Unit 5 of the Serra plant of the Sao Paulo Tramway Light and
Power Company, Ltd., is a double-overhung impulse turbine of
which the nominal rating is 84,000 hp at 680 m (2230 ft) net
head and 360 rpm; on test it can produce 92,300 hp at 674 m
(2210 ft) head. It is fed by a penstock line approximately 1600
m (5250 ft) long with three diameters; 1565 mm (61.61 in.),
1463 mm (58.60 in.), and 1346 mm (52.99 in.). Fig. 1 shows in
sketch form the upper end of this penstock and the location of
the salt-velocity injection station and the three electrode sets
used for testing the unit.

664 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
The injection station consisted of four 2-in. pop valves set in
from the pipe wall 38 per cent of the pipe radius. Elementary
electrodes were placed at the side of two of the pop valves to determine
the time of the salt-solution injection.
Each electrode set consisted of two pairs of electrodes at right
angles to each other, extending across the penstock to within
V2 in. of the penstock walls. These electrodes were bowed to
have a center spacing 6.5 times that of the electrode ends.
The volume between the axes of the electrodes was determined
from four measurements of diameter in each pipe, approximately
one measurement for each 5 ft of penstock, and tape measurements
of the lengths, due allowance being made in the volume for
the bump joints and rivet heads, from the penstock manufacturer’s
drawings.
In testing the turbine, a total of 24 runs with an average of 5.6
salt shots per run was made. For each run six different determinations
of the discharge were obtained, i.e., from the injection
station X to electrode sets A, B, or C; from electrode set A to B,
or to C; and from electrode set B to set C.
Table 1 shows the number of salt shots for each run, the time
of salt passage for each run and electrode combination, the resulting
discharges, and the variations of these discharges from
those obtained from the combination of electrode sets A and C.
These variations have been analyzed by computing the probable
variation of any one run for the several sections with the
formula
where r = probable error, per cent
v = variation, per cent
n = number of runs
The probable variations of one run, as percentages, are as follows:
Section...........X — A X — B X — C A — B B — C
r................ ±0.47 ±0.22 ±0.10 ±0.12 ±0.12
This shows a high degree of consistency of the discharge measurements,
the greatest probable deviation of one run being less
than 0.5 per cent for section X — A where the testing setup is the
poorest. Examination of the variations shows no tendency for
consistent positive or negative variations. In the case of section
X — A, this tendency which is there the greatest, averages
—0.07 per cent for the 24 runs, but is meaningless, since the
probable deviation of the average is greater, being ±0.10 per
cent.
A further indication of the small variation in the discharge is
obtained from the calibration of the venturi restriction based on
discharges determined in section A — C. The least-squares
determination of a straight line fitting the observations gives K
— 7.807 + 0.0105 Q in the formula Q = K y/H, where
Q = discharge, cu m per sec
H = venturi differential head, m of water
The variations of the observations from this line show a probable
variation of any one point of ±0.0148 or say ±0.19 per cent.
Considering that this variation includes the errors of measuring
the differential head by a U tube containing bromoform, as well
as the errors of the discharge measurement, it is in line with the
variations found in the discharge determinations of section
X — C, A — B, ox B — C.
S e r r a U n i t s 3 a n d 7
Unit 7 of the Serra plant is essentially a duplicate of unit 5
previously described. Unit 3 is similar, but with slightly smaller
penstock and nozzle diameters, having a rated output of 72,800
hp and an actual maximum of 78,100 hp. The penstocks for
units 7 and 3, especially the upper end, and the salt-velocity
setups as applied there, were similar to that for unit 5.
When unit 3 was installed and tested in 1936, its penstock was
temporarily connected to the venturi restriction and butterfly
valve intended for unit 7. Sixteen calibration runs were obtained
giving an average of K in the equation Q = K \ / H of 7.447 ±
0.006. The probable deviation of one run was r = ±0.32 per
cent.
Later in 1939, the penstock of unit 3 was disconnected from
this venturi and the penstock of unit 7 connected. The venturi
was again calibrated during the tests of unit 7. Twenty-two
calibration runs were obtained, giving an average value of K
of 7.430 ± 0.005. The probable deviation of any one run was
r = ±0.28 per cent.
In either case, the salt-velocity setup was similar to that for
unit 5. The discharges were determined from electrode sets A
and C. The differential heads were measured by U tubes containing
bromoform. The data of these two calibrations are given
in Table 2 and are plotted in Fig. 2.
In the case of unit 3, the location of the injection station was

DODKIN—FIELD CHECKS OF TH E SALT-VELOCITY METHOD 665
only 11 m, 7 pipe diameters, below the butterfly valve and venturi
restriction, as compared with 24 m, 15.5 pipe diameters, for unit
5. The distorted velocity distribution resulting from the venturi
and butterfly valve upset the precision of the salt-velocity discharges
determined from the injection station to the electrodes.
Such determinations differed from those from electrode sets A to
C by as much as —2 per cent for X to A decreasing to —0.35
per cent for X to C. The results between electrode sets A — B
and B — C, when compared with those between electrode sets
A — C, were satisfactory, the variations ranging from 0.59 per
cent to zero, with a probable deviation for any one run of r =
±0.21 per cent.
The foregoing shows one point in the application of the saltvelocity
method which must be emphasized. Where a high degree
of accuracy is desired, the timing of the passage of the salt
cloud in the waterway must be done between sets of electrodes
extending across the entire waterway, not from the injection
station to one electrode set, although with care in location of the
injection station good results may be obtained thus, as shown by
the tests on unit 5.
Parahyba U nit 2
Unit 2 of the Parahyba plant of the Brazilian Hydro-Electric
Company, Ltd., is a vertical reaction-type turbine, rated at
28,000 hp at 28.66 m (94 ft) net head or 32,500 hp at 31.7 m
(104 ft) head and at 125 rpm. On test it exactly equals its rating.
Two complete efficiency tests were made on this unit December
19 and 21, 1927. On the first test, the discharges were determined
F i g . 3 T e s t i n g L a y o u t a t P a r a h y b a U n i t 2

666 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
G — GUID E-VAN E O P EN IN G , P E R C EN T
F ig . 4 T e s t R e s u l t s ; P a r a h y b a U n it 2
by the Gibson method applied differentially (10 runs). On the
second test they were determined by the salt-velocity method
(15 runs). Conditions of head, etc., were essentially identical
on the two days.
Fig. 3 shows schematically the testing layout. Expressing the
lengths in terms of the pipe diameter (average for the part concerned)
the salt-velocity mixing length, introduction to upper
electrode set, was 2.7 and the test length between electrode sets
was 6.82. The length between the Gibson piezometers was 2.85.
For the salt velocity, 12 pop valves were used at the stop-log
plane. This gave one valve for each 40.5 sq'ft of cross section
at the injection station or one valve for each 24 sq ft of conduit
section at the upper electrode set. The path from the highest
pop valves to the upper electrode set was about 17 per cent longer
than from lowest pop valves.
It was recognized from the start that, because of the short
lengths relative to the diameter, and because the plane of the
pop valves was not parallel to that of the upper electrodes, the
testing setup was not ideal. Consequently, precautions were
taken to insure good results with both methods. Great care
was taken to adjust all pop valves for tightness up to 15 lb per sq
in., and to have these open uniformly and at the same pressure.
The electrodes were carefully shaped to have a center spacing 6.5
times the end spacing. The large electrode area and the natural
conductivity of the water required an applied voltage of about
10 v. Consequently, large wires with carefully made joints were
used in the electrode wiring.
For the Gibson method, all piping joints were carefully made
for tightness. Also, during the performance of tests, the piping
was flushed through completely between each run to eliminate
any air and to insure that the temperature of the water in the
piping was essentially that of the penstock water.
Table 3 presents the data obtained on these two tests. Fig. 4
shows the curve of discharge against guide-vane opening after
transferring to a uniform net head of 28.66 m. In order to present
graphically and clearly the variations between the discharge
determinations by the two methods, and in an attem pt to rectify
the data for analysis by the method of least squares, each observed
discharge at 28.66 m head has been divided by the corresponding
guide-vane opening expressed in per cent. These quotients
are likewise plotted in Fig. 4.
Examination of the plotting of these quotients shows that they
may be approximately represented by straight lines from a gate
opening of 53 per cent up to 100 per cent. The least-squares determination
of the two equations of such lines gives
and
Q/G = 1.414 — 0.00390 G for the Gibson data
Q/G = 1.406 — 0.00385 G for the salt-velocity data
where
and
Q — discharge at 28.66 m head in cu m per sec
G = guide-vane opening in per cent.
The equation determined by the Gibson data in comparison
with that determined by the salt-velocity data gives 0.42 per cent
greater discharge at 53 per cent guide-vane opening (lower limit
of applicability) and 0.29 per cent greater discharge at full guidevane
opening. The computed values of the probable variation of
one run from the equations are r — =*=0.48 per cent for the Gibson
method and r = ±0.22 per cent for the salt-velocity method.
P a r a h y b a U n i t 4
Unit 4 of the Parahyba plant is a vertical reaction-type turbine

DODKIN—FIELD CHECKS OF TH E SALT-VELOCITY METHOD 667
rated at 54,000 hp at 31.7 m (104 ft) net head and 115.4 rpm,
with a discharge at maximum gate of over 5500 cfs; on test it
can produce 57,190 hp at 31.7 m head.
A complete efficiency test of this unit was made in November,
1938. On this test, 28 runs were made using the salt-velocity
method and, for 17 of these runs, the Gibson method was used at
the end of the 10-min test run.
Fig. 5 shows schematically the testing layout which is generally
similar to that for unit 2, except the pop-valve arrangement and
the forebay water elevation at the time of the test. The lengths
in terms of the pipe diameter are, salt-velocity mixing length 1.6,
salt-velocity test length 6.4, and Gibson piezometer length 4.2.
For the salt-velocity method, 20 pop valves were used in a
plane crossing the stop-log groove and closely parallel to the
plane of the upper electrodes, thus making the paths from all
valves of the injection station to the electrodes essentially equal.
This gave one valve for each 35.5 sq ft of vertical cross section
at the injection station or one valve for each 20.7 sq ft of conduit
section at the upper electrode set. The same general precautions
for insuring accuracy were followed with both methods as were
used for unit 2.
Table 4 presents the data of this test. Fig. 6 shows the curve
of discharge against guide-vane opening after transferring to a
uniform net head of 31.7 m. As for unit 2, in order to present
more clearly the variations of the discharge determinations and
in an attem pt to rectify the data for analysis by the method of
least squares, each observed discharge at 31.7 m head has been
divided by the corresponding guide-vane opening in millimeters.
These quotients are likewise plotted in Fig. 6.
Examination of the plotting of these quotients shows that they
may be approximately represented by straight lines over the
range covered jointly by the Gibson and salt-velocity observations,
that is, from 46 to 89 per cent guide-vane opening. Thus
using all Gibson points and omitting the salt-velocity determinations
of runs 1, 4 to 9, and 28, as being outside the range of joint
observations, the least-squares determination of the two equations
of such lines gives
and
Q/G = 0.5991 — 0.0004455 G for the Gibson data
Q/G = 0.6039 — 0.0004594 G for the salt-velocity data
where
and
Q = discharge at 31.7 m head in cu m per sec
G = guide-vane opening, mm.
The equation determined by the Gibson data in comparison
with that determined by the salt-velocity data gives 0.42 per cent
smaller discharge at 180 mm guide-vane opening (lower limit of
applicability) and identical discharge at 345 mm guide-vane
opening (upper limit of applicability). The computed values of
the probable variation of one run from the equations are r =
=*=0.0024 or say =*=0.5 per cent for the Gibson method and r
= ±0.0031 or say =*=0.6 per cent for the salt-velocity method.
S o r o c a b a U n i t 4
Unit 4 of the Sorocaba plant of the Sao Paulo Electric Company,
Ltd., is a horizontal reaction-type turbine rated at 25,000
hp at 195 m (640 ft) net head at 600 rpm; on test it can produce
26,450 hp at 195 m head. Water is supplied to the unit by a
penstock line about 656 m (2150 ft) long, having diameters of

668 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
1850 mm (72.83 in.), 1750 mm (68.90 in.), and 1650 mm (64.96
in.).
Fig. 7 shows schematically the layout of the plant from the
lower end of the canal to the powerhouse, together with the profile
of unit 4 penstock.
A complete turbine-efficiency test of the unit was made in
January, 1926. Fourteen runs, measuring the discharge with the
salt-velocity method were made. As shown in Fig. 7, the measurements
of discharge were made between two sets of electrodes
of the customary bowed type. The salt-injection station consisted
of 4 pop valves set in from the pipe wall 39 per cent of the
pipe radius and was generally similar to the Serra layout in Fig. 1.
F i g . 7 T e s t i n g L a y o u t ; S o r o c a b a U n i t 4

DODKIN—FIELD CHECKS OF TH E SALT-VELOCITY METHOD 669
The results of the test appeared entirely satisfactory, check
points at identical gate openings showing a maximum difference
of 0.2 per cent. However, the venturi meter, having full and
throat diameters of 1850 and 1269 mm, respectively, showed a
coefficient about 4.5 per cent lower than the usually expected
value of 0.985.
As these tests were made in the early days of the companies’
experience with the salt-velocity method, a volumetric check of
the discharge was made as this was possible with the plant layout.
The volume of the canal forebay, and surge tank between three
elevations, was determined by measurement of the areas. With
all inflow shut off except a very small but measured inflow from
a spring and no other outflow, this known volume was discharged
through penstock 4 by bringing the unit from complete shutdown
to a predetermined gate setting in about 2 min, holding this load
constant for as long a period as the available water permitted,
and shutting down the unit completely in about l 1 A min. The
rate of discharge during the steady-load portion of the run was
determined by deducting the percentage of water used during
starting and stopping, determined on the basis of the previous
constant-load tests from gate openings observed every 10 sec
during changes of load. The water used in starting and stopping
amounted to 16 per cent of the total on one test and 7 per cent
on the other two tests. During the steady-load portion of each
run, salt-velocity measurements were made.
T A B LE 5 V O L U M E T R IC C H E C K O F SALT V E L O C IT Y ;
SOROCABA U N IT 4
Item T est 1 T est 2 T est 3
804.688 804.584 804.638
Final level, m ................................................. 803.858 801.645 801.535
Average area above 803.57, sq m .......... 6130 5640 5700
D raft above 803.57, m ............................... 0.830 1.014 1.068
Average area below 803.57, sq m ........... 3950 3940
D raft below 803.57, m ............................... 1.925 2.035
Volume, cu m ................................................ 5088 13323 14106
Inflow from spring, cu m ........................... 20 27 22
Total w ater used on test, cu m ............... 5108 13350 14128
W ater used on steady load, per c e n t.. . 84.15 93.20 92.70
W ater used on steady load, cu m .......... 4298 12442 13097
Tim e of steady load, sec.......... ................. 720 2100 1700
Average discharge— volum etric, cu m
per sec.......................................................... 5.969 5.925 7.704
Average tim e salt velocity, sec ............... 164.9 167.0 127.9
Average discharge— salt velocity, cu m
per sec......................................................... 5.967 5.892 7.694
Difference salt velocity— volum etric,
per cen t....................................................... —0.03 —0 .5 6 —0 .1 3
Table 5 presents the summary of these tests. It will be seen
that the discharges determined volumetrically and from the saltvelocity
method agree within 0.25 per cent average for the three
tests. It is considered that the penstock volume for the saltvelocity
and the canal volume for the volumetric check as determined
had each a precision of about =*=0.3 per cent. The test
agreement is therefore as good as can be expected. At that time
no greater refinement in testing results was attempted.
G e n e r a l C o m m e n t o n T e s t s
From the period of obtaining the rights to use the salt-velocity
and Gibson methods to date, all units of the principal plants of
the companies associated under the Brazilian Traction, Light and
Power Company, Ltd., have been tested by one or the other of
these methods, with the exception of three units which are duplicates
of machines tested. In some cases, the tests have been made
indirectly by using the Gibson or salt-velocity method to calibrate
a venturi restriction, using this calibrated restriction as a measuring
device for several units. Eleven units have been directly
tested with each method, a total of 16 tests with the salt-velocity
and 15 tests with the Gibson method having been made. All of
these tests have been satisfactory and, where doubts have occasionally
been expressed, repeat tests and checks by other
methods have confirmed the accuracy of the tests.
From the experience gained on these tests, the author suggests
the following, but by no means comprehensive, precautions regarding
the use of the two methods:
Salt-Velocity Method
1 If best results are desired, the timing of the passage of the
salt cloud must be done between two sets of electrodes which
represent the entire cross section of the waterway.
2 An adequate mixing length between the injection valves
for the brine and the first set of electrodes is imperative for best
results. The author does not feel competent to place a definite
limit on this length, 'but suggests a lower limit of a length at least
equal to the mean transverse dimension of the waterway and
preferably at least twice this length. It appears also that the
number of pop valves used should depend on this length rather
than on the cross-sectional area of the waterway.
Gibson Method
1 The author has found that measuring the leakage with
fully closed guide vanes may, at times, be the most troublesome
feature of this method. In two cases it appeared conclusively that
the leakage through the guide vanes did not vary as the square
root of the head on the turbine, thus making it necessary to measure
the leakage with the full head. No explanation of the discrepancies
has been found, but they may be due, at least in part,
to the variation in mechanical deflection of the guide vanes with
varying applied pressure.
2 Besides the necessity of absolutely tight piezometer piping,
the author considers it necessary to provide adequate provision
for flushing the piping between tests to remove air and to maintain
the penstock water temperatures if best results are to be
obtained.
A c k n o w l e d g m e n t s
The author wishes to make acknowledgment to A. W. K.
Billings, member of this Society, and vice-president of the
Brazilian Traction, Light and Power Company, Ltd., for permission
to use the companies’ data, to C. P. Conrad and L. J. Hooper,
also a member of this Society, who made the Sorocaba tests and
to A. Santos, Jr., who made some of the many tests carried out
by the companies and assisted in the performance of the others.
D iscussion
M . A. M a s o n .4 The writer had the opportunity, while studying
in France as a Freeman scholar, to make an investigation of
the possibilities of use of the salt-velocity method in open channels.
It was found that previous studies had not resulted in the
formulation of any logical conception of the physics of the method
which would permit a definition of its limitations. An attem pt
was made, therefore, to define the limits of applicability of the
method by a study of the basic principles governing its operation.
The conclusion was reached that the chief requirement for accurate
results is the attainment of proper mixing between the
water and injected solution before the injected solution reaches
the measuring section.6 It is gratifying that the author arrives
at an identical conclusion for the case of flow in pipes.
Concerning the magnitude of the “mixing length” required for
accurate results, it was concluded from the writer’s studies that
this length was probably a function of the velocity of flow, the
injection pattern, the form of the channel, the nature of the walls,
and some time period. The author presents no evidence to
* Engineer, Beach Erosion Board, W ar D epartm ent, W ashington,
D. C. Jun. A.S.M.E.
• “ Contribution & l’fitu d e de la meaure des D ebits d ’Eau par la
Mfithode Allen,” by M. A. Mason, Thesis, University of Grenoble,
France; also R6vue G6n6raXe de VHydraulique, No. 28, July-Aug., 1939.

670 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
justify the elimination of any of these factors and his conclusion
that the “good mixing length” may be specified in terms of only a
linear dimension of the conduit system does not seem to be
justified.
Professor Hooper, in his paper on salt-velocity measurements
of low-velocity flow8 employs mean velocity of flow as a measure
of the mixing of injected solution and water. It is probable
that both criteria suggested are involved in the true criterion for
good mixing and that neither is sufficient in itself.
Although no concrete evidence of the effect of electrode design
is presented in this paper, the writer concurs in the first conclusion
to the effect that the electrodes must suitably cover the
entire cross section of the waterway. An analysis leading to this
conclusion was included in the writer’s paper.6
E. B. S t r o w g e r .7 The author indicates that he has obtained
a satisfactory agreement in the results of tests of two units by the
salt-velocity method with those of the differential application
of the Gibson method. The writer wishes to submit the results
of other tests which show the consistent accuracy of the
Gibson method.
Perhaps it should be pointed out that the Gibson method
makes use of pressure-time diagrams of either the simple or the
differential type. With simple diagrams, the changes of pressure
at one piezometer section in the conduit are recorded while
with differential diagrams the difference between the changes of
pressure at two piezometer sections in the conduit are recorded.
Indirect checks of the Gibson method can be made, therefore,
when both simple and differential diagrams are used; in tests
where two sets of simple diagrams are made, each utilizing a
different length of penstock; in tests where two differential applications
are made, each utilizing different lengths of pipe;
and/or when tests are made using other comparisons.
Some of the material which will be presented herein has been
published before but is being included for the purpose of assembling
as much of this information in one place as is possible at
this time.
Fig. 8 of this discussion shows the cross section of a 70,000-
hp hydroelectric unit operating under a head of 217.5 ft. It also
shows the piezometer sections used in testing this unit by the
Gibson method. Simple diagrams were made on the unit,
utilizing approximately 600 ft of penstock, extending from the
* “ Salt-Velocity M easurements at Low Velocities in Pipes,” by
L. J. Hooper, published on page 651 of this issue of the Transactions.
7 Hydraulic Engineer, The Niagara Falls Power Company, Buffalo,
N. Y. Mem. A.S.M.E.
DISCHARGE, CFS
F i g . 9 P o w e r -D isc h a r g b C u r v e O b t a in e d b y G ib s o n T e s t o n
U n it S h o w n i n F i g . 8, W it h T e s t P o in t s o f B o t h S im p l e a n d
D if f e r e n t ia l D ia g r a m s
forebay to the lower piezometer section and differential diagrams
were made utilizing about 87 ft of the penstock as shown. The
agreement between the two tests is indicated by the two sets of
test points which fall on or close to the same power-discharge
curve as shown in Fig. 9. The average divergence of the “differential
points” from the line established by the “simple points” is
within +0.2 per cent and the agreement between the two tests is
accordingly within 0.2 per cent.
Fig. 10 shows the plan and cross section of the penstock of a
14,000-hp unit operating under a head of 96 ft. On this unit two
tests were made by the Gibson method, one using simple dia-

DODKIN—FIELD CHECKS OF TH E SALT-VELOCITY METHOD 671
CRO SS SECTION
F i g . 10 P l a n a n d C r o s s S e c t i o n o f P e n s t o c k o f 14,000 U n i t ,
S h o w i n g P i e z o m e t e r S e c t i o n s f o b G i b s o n M e t h o d o f T e s t i n g
F i g . 13 P o w e r D is c h a r g e C u r v e f o r 1 0 ,0 0 0 -H p U n i t ;
H e a d 212 F t
C r o s s
(T est No. 1, Oct. 4, 1921, represented by points m arked O m ade w ith Gibson
a pparatus No. 5 attached ju st below 90-deg elbow. T est No. 2, Dec. 11, 1921,
represented by points m arked A m ade w ith Gibson a pparatus No. 1 a t­
tached upstream from 90-deg elbow. In te st No. 1, F = 4.010, R = 2.35,
and K = 118.50. In te st No. 2, F - 3.672, R = 1.528, K - 89.35.)
F i g . 12 S e c t i o n T h r o u g h P e n s t o c k o f 55,000 H p T u r b i n e ,
S h o w i n g L o c a t i o n o f P i e z o m e t e r s f o r G i b s o n A p p a r a t u s
grams and the other using differential diagrams. For the simple
diagrams, all of the pipe was utilized extending from the forebay
to piezometer section B and, for the differential diagrams, the
section between piezometer sections A and B was utilized. The
two sets of test points are plotted in Fig. 11, and determine one
curve as shown.
In Fig. 13 is shown the power-discharge curve of a 10,000-hp
unit operating under a 212-ft head as determined by two sets of
test points, each made with the simple diagram application of
the Gibson method but with different piezometer locations.
One piezometer section was located upstream from a 9-ft elbow
and the other just downstream from the elbow. The two sets

672 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
F i o . 1 4 G a t e O p e n i n g D i s c h a r g e a n d G a t e O p e n i n g P o w e r C u r v e s f o r 5 5 ,0 0 0 H p T u r b i n e ; N e t H e a d 3 1 2 F t ; b y G i b s o n
M e t h o d , D e c . 1 0 , 1 9 2 2
(In t e s t N o . 1 O , F - 1 .7 4 2 , R - 1 .8 0 2 , K - 9 9 .4 5 . In t e s t N o . 2 A , F - 1 .3 8 5 , R - 1 .5 1 5 , K = 8 8 .9 8 .)
F i g . 1 5 P i e z o m e t e r S e c t i o n s f o r T w o S e t s o f D i f f e r e n t i a l
D i a g r a m s , G i b s o n M e t h o d o f T e s t i n g , 1 0 , 0 0 0 - H p U n i t
(L e n g th A to B , 4 2 .1 0 f t. L e n g th A t o C , 7 9 .5 4 f t.)
of test points indicate that a common power-discharge curve may
readily be drawn.
Another example of two sets of simple diagrams taken on the
same penstock is illustrated by the test on a 55,000-hp unit operating
under a head of 312 ft, as shown by Figs. 12 and 14. With
one set of diagrams approximately 310 ft of penstock was utilized
and with the other set approximately 367 ft was used. The
diagrams were made simultaneously, i.e., at one closure of the
turbine gates, two diagrams were made. The two sets of test
D i s c h a b .s e ~ c f s .
F i g . 1 6 P o w e r - D i s c h a r g e C u r v e F r o m G i b s o n T e s t o n 1 0 ,0 0 0
H p U n i t ( P e n s t o c k S h o w n i n F i g . 1 5 )
points shown by triangles in the one case and circles in the other
were determined by two independent testing organizations using
different apparatus and working independently.
The penstock shown in Fig. 15 was used for making two sets of
differential diagrams, one set using 42.10 ft of length and the

DODKIN—FIELD CHECKS OF TH E SALT-VELOCITY METHOD 673
T A B L E 6
C O M PA R ISO N O F R E SU L T S O F T E S T S M A D E IN E U R O P E B Y G IB SO N
M E T H O D W IT H T H O SE O F O T H E R M E T H O D S
No. D ate Place C onducted by
1 O ctober, 1928 Heidenheim , G erm any J . M . Voith
2 July, 1929 W alchensee, G erm any P aul V olkhardt
3 A ugust, 1929 E itting, G erm any P aul V olkhardt
4 R oyal School of Engi- E tto re Scimemi
neers, Padova, Italy
5 1930 Technischen Hochschule, H ans Deckel
MOnchen, G erm any
6 1930 BrQnnenmuhle in Heid- H ans Deckel
enheim
Results
M ean variation from weir
m easurem ent, 1 .9 per
cent
M ean variation from current
m eter m easurem ent, — 0.3
per cent
M ean variation from current
m eter m easurem ent, + 0 .4
per cent
M ean variation from volum
etric m easurem ent, — 0.2
per cent_
M ean variation from volum
etric m easurem ent, — 0.3
per cent
M ean variation from weir
m easurem ent, — 0.4 per
cent
DISCHARGE, CFS
DISCHARGE , CFS
Fio. 17 C o m p a ris o n o f R e s u l t s , G ib s o n T e s t a n d A l l e n T e s t Fiq. 18 C o m p a r is o n o f R e s u l t s , G ib s o n T e s t a n d A l l e n T e s t
o n 59,000-Hp U n i t ( U n i t N o . 1) N e t H e a d = 196 Ft
o n 59,000-Hp U n i t ( U n i t N o . 2) N e t H e a d =■ 196 Ft
F i o . 19 C o m p a r is o n o f R e s u l t s , G ib s o n a n d A l l e n T e s t s , 1900
H p U n i t , N e t H e a d = 8 9 F t
F iq . 20
D ata o n V e n t u r i S e c t io n s , R a t e d in t h e F ie l d b y t h e
G ib s o n M e t h o d

674 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
other set a length of 79.54 ft. The power-discharge curve with
the two sets of points is shown in Fig. 16.
Figs. 17, 18, and 19 show the results of testing three units by
both the Allen and Gibson methods. The close agreement between
the two measurements in each case is apparent.
Fig. 20 shows the coefficient curves of two 7 X 10-ft ven­
DISCHARGE IN HUNDREDS OF CFS
F ig . 21 ( L e f t ) R a t i n g C u r v e M a d e b y G ib s o n M e t h o d o f a 120
X 8 4 -In . V e n t u r i ( M e t e r N o . 1)
F ig . 22 ( R ig h t) R a t i n g C u r v e M a d e b y G ib s o n M e t h o d o f a 120
X 8 4 -I n . V e n t u r i ( M e t e r N o . 2)
turi meters built in the two penstocks of a hydroelectric station
and tested by the Gibson method when the acceptance tests of
the units were made. It shows the test points plotted as velocity
against venturi coefficient. Figs. 21 and 22 show the test
points plotted as discharge against manometer deflection on log
paper indicating that, with this plotting, the relation between deflection
and discharge is a straight line. The slope of the line is 2
and accordingly the discharge varies as the square root of the
mercury deflection. This means that the coefficient for these
venturis is constant for the range of velocity covered by the test.
Table 6 of this discussion presents a list of comparative test
results using the Gibson method in conjunction with other
methods of water measurement as carried out at various places
in Europe. The results of these comparisons have been published
from time to time in the technical press. With the exception
of the first tests at Heidenheim, which were not very good, all
these results are within the practical limits of precision in the
measurement of flowing water.
The author states he has found at times that measuring the
leakage of the turbine with closed guide vanes was the most
troublesome feature of the Gibson method and mentions two
cases where the leakage did not vary as the square root of the
head on the turbine. He suggests that the discrepancy, in part
at least, was due to the mechanical deflection of the guide vanes
with varying pressure. Since in many cases the turbine gate
leakage must be determined under heads which are lower than
the normal value and the experimental result must be stepped
up to the full head, care must be exercised to see that proper
adjustment of the orifice area is made in case deflection of the
guide vane stems takes place. If this is done, the writer believes
that the author will not consider the measurement as troublesome.
Gate-leakage determinations usually can be made at
several values of reduced head, especially if the fall in the pressure
method of measuring the leakage is used, and the results
may be plotted in the form of a leakage-head curve on log paper.
If the orifice area changes with head, the relation so obtained is a
curve rather than a straight line and may be extrapolated to the
full head. The area correction is made in the process of extrapolating
the curve.
Another method of determining whether an area correction is
needed in stepping up the leakage to the full head is to compare
the “squeeze” deflection of the wicket gates under full operating
oil pressure with the calculated deflection of the wickets under
full head conditions. If the squeeze deflection is greater than
the deflection due to casing pressure, no area correction is necessary.
An example of such a computation is shown in Fig. 23 for
one of the units in a hydroelectric plant in New York State. At
this plant the servometer piston travel is limited by a stop which
is so placed that the wicket gates are closed with no squeeze when
the wheel case is empty. The computed deflection of the guide
vanes with full casing pressure was 0.093 in. resulting in an orifice
area of 0.233 sq ft and a computed leakage of 28.6 cfs. The actual
measured leakage per unit under full head as determined by a
tailrace weir was 29.5 cfs. Other cases of leakage determination
where the guide vane deflection is a factor in the determination
could be cited, although with the method used in assembling
and adjusting the gate mechanism of modem hydroelectric units,
such deflection is seldom found to be appreciable.
It should be noted also that, as a practical matter, extreme care
in measuring the leakage is not as a rule necessary. A relatively
large error in the determination of the turbine-gate-leakage
quantity affects the turbine efficiency by an insignificant amount.
For example, on a recent test, a possible error in the leakage
quantity of as much as 20 per cent of the total leakage represented
less than 0.1 per cent in turbine efficiency at the point of
maximum efficiency.

DODKIN—FIELD CHECKS OF THE SALT-VELOCITY METHOD 675
E d w i n A. T a y l o r .8 The results of three check tests at Sorocaba
show that the salt velocity varied from volumetric by 0.03,
0.56, and 0.13 per cent, respectively. These results have a maximum
spread for any one run of 0.5 per cent and an average difference
between the two methods of 0.25 per cent.
The results of two check tests at Parahyba show that, for the
wide range of the tests, salt-velocity and Gibson methods check
within less than 0.1 per cent. For each unit, the two curves of
discharge on gate opening coincide at full gate and differ only
slightly at half gate.
8 Consulting Engineer, W orcester, Mass.
A check test at Serra No. 5, comparing discharge results,
using three different salt-velocity test sections, shows a maximum
spread of less than 1 per cent, 0.5 per cent plus and minus. For
the whole test, the average difference was less than 0.1 per cent.
These comparative tests covered 24 runs and 130 salt-velocity
shots. Similar check tests, with equally close comparative results,
were made at Serra Nos. 3 and 7.
The writer has made a number of pump- and turbine-efficiency
tests, using the salt-velocity method of water measurement.
Several times the results by salt velocity have been checked by
other engineers, using other methods of water measurement,

676 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
and frequently the salt-velocity results have been checked by the
same method, using different test apparatus and test sections at
the same plant.
Tables 7 and 8 show the results of fourteen such check tests
on salt velocity, two by current meter, two by pitometer, two by
Gibson, two by volumetric, and six by interchecking with various
salt-velocity test sections.
Details of tests, apparatus, and procedure for the check tests in
Tables 7 and 8 are omitted. The engineers conducting these
tests were experienced in their lines and the tests were made in
accordance with the latest accepted practice in each method.
A summary of the comparative results, shown in Table 7 of this
discussion, indicates a close agreement between salt velocity and
current meter, once in Germany and once in Pennsylvania. A
comparison of the discharge values by pitometer in Germany and
in New York and by Gibson in Quebec and in New York shows a
close agreement with the salt-velocity values for discharge at the
same plants.
A comparison of salt velocity with the five volumetric check
runs in Table 7 shows the results have an average spread for all
runs of 0.75 per cent and an average weighted difference for the
five runs of 0.6 per cent.
While the values for average difference for the Pacific Coast
volumetric checks are plus 0.6 per cent, the Brazilian volumetric
checks show an average difference of minus 0.25 per cent.
The average mean difference, weighted for signs, for the eight
check tests in Table 7 is +0.05 per cent.
In analyzing and comparing the results in Table 7, the spread
and the deviation for runs should be given more weight than the
average difference in values for the whole test. This average
difference may mean very little alone. For instance, there
could be a spread or run deviation in discharge values of 5 or
even 10 per cent and the mean difference between methods for the
whole test could still be zero. In that case, attaching much importance
to the mean difference for the whole test becomes very
significant of the accuracy of the method. This is important in
field tests for pump and turbine efficiency because, in computing
final results, curves are drawn to fit the test points and these
curves do graphically what the average differences in Table 7 do
arithmetically.
Table 8 of this discussion shows the results of six check tests at
four different power plants, using various salt-velocity test sections.
These test sections varied in length from 300 ft to 3.5
miles. The average spread for all runs shown in Table 8 was
0.33 per cent. The average mean difference, weighted for
signs, for the six check tests shown in Table 8 is minus 0.06 per
cent.
The mean difference for all the check tests shown in both
Tables 7 and 8, which involve over 100 salt-velocity runs and 500
shots, is zero.'
A u t h o r ’s C l o s u r e
Messrs. Strowger and Taylor have presented additional data
showing verifications of the salt-velocity and Gibson methods.
The assembly of such data is valuable in appraising the degree
of accuracy which may be expected by these methods. Further
testing results, where the agreement was less satisfactory, would
also be valuable. Undoubtedly, attempts to extend the field of
these methods have been responsible for the majority of tests
where agreement was relatively poor. A record of these attempts,
and of the experience gained in overcoming difficulties, is desirable.
To present one such case the following is submitted: Prior
to the tests of Parahyba unit 4 described in the paper, preliminary
tests were made on this unit with both salt-velocity and Gibson
methods. The electrode setup for the salt velocity was the same
as in the later tests. The injection station consisted, however,
of only 12 pop valves placed in the vertical plane of the stop-log
grooves. The form of the resulting curves was decidedly poor, each
curve showing 3 peaks resulting from the unequal paths from the
3 horizontal rows of 4 pop valves each. The results differed from
the final results given in the paper by an amount varying from 0
per cent at 180 mm guide-vane opening to +1.6 per cent at 345-
mm guide-vane opening. The use of the salt-velocity method
in this manner was an obvious but unwitting abuse. Notwithstanding,
there is comfort in the relative closeness of agreement
with the later tests made with an improved layout. The preliminary
Gibson tests agree very closely with the final ones reported
in the paper.
The suggestions given in the paper regarding mixing length
resulted principally from thought on conditions of testing at the
Parahyba plant. Mr. Mason, in his study6 and in his discussion
of this paper, and Professor Hooper, in his paper® have contributed
further to this question. The determination of conditions providing
a satisfactory mixing and distribution of the salt solution
in the water channel between the injection valves and the upper
electrodes would be a suitable laboratory study. It is apparent,
however, that a requirement relating the number of injection
valves to the cross-sectional area of the waterway is entirely inadequate
without further stipulations. These probably should
include factors for the channel dimensions, form, roughness, and
possibly for the velocities expected. For ordinary commercial
testing in hydroelectric plants the low velocity limits brought out
by Professor Hooper are rarely, if ever, encountered.

The V iscosity of S uperheated Steam
B y G. A. HAWKINS,1 H. L. SOLBERG,2 a n d A. A. POTTER,3 LAFAYETTE, IND.
After discussing the disagreement existing in viscosity
data for superheated steam as reported by recent investigators,
the authors report the results of a new determ ination
of viscosity. A nickel capillary 103.2 ft long was used
in this investigation. The differential head across the
capillary was determined by radiographing the mercury
levels in a steel-tube m anom eter. Data are reported for
60 calibration tests and for 61 tests on steam at pressures
up to 1810 lb per sq in. abs and 1000 F. All o f the tests were
run under steady-flow conditions for a m inim um period of
I hr. Plank’s equation has been modified to express the
results over the entire range of pressures and tem peratures
covered by this investigation.
As a result of additional calibration tests made to determine
the effect o f curvature on the capillary constant, the
results given in this paper supersede those presented at
the Annual M eeting of the Society.
IN THE year 1925, Speyerer (l)4 measured the viscosity of
superheated steam by means of a capillary tube at pressures
from 1 to 10 atm and at a maximum temperature of 658 F.
In 1934, Schiller (2) published data on the viscosity of superheated
steam as determined from the flow of steam through a
nozzle. He obtained the relation for calculating the viscosity of
Bteam in terms of that of water, by observing the velocity at
which the discharge coefficient of a nozzle shows an abrupt increase
for water and steam and then equating the two Reynolds
numbers. Schiller’s curves show one test point at 30 atm, four
test points at 25 atm and other points at lower pressures with a
maximum temperature of about 540 F. In 1935, the authors (3)
published data on the viscosity of saturated and compressed
water at temperatures up to the critical and on superheated
steam at pressures up to 3500 lb per sq in. abs and temperatures
up to 983 F. A modified form of the Lawaczeck (4) viscometer
was used in which the time of fall of a metal cylinder in an accurately
bored stainless-steel tube was recorded automatically.
The results obtained by the authors (3) were in good agreement
with those of Schiller and Speyerer.
In 1934, Schougayew (5) published data obtained by measuring
the viscosity of steam in a Rankine capillary at pressures up to
93 atm and temperatures up to 400 C. He concluded that, within
the limits of accuracy of his apparatus, viscosity was practically
independent of pressure. In a discussion of the author’s paper,
Schougayew (6) presented a curve, showing the difference between
the results for the viscosity of saturated steam, as plotted from
the author’s data, and similar results obtained in Russia by
Schougayew and Sorokin.
1 Associate Professor of Mechanical Engineering, Purdue University.
Mem. A.S.M.E.
2 Professor of Mechanical Engineering, Purdue University. Mem.
A.S.M.E.
3 Dean, Schools of Engineering, Purdue University, and Director
of the Engineering Experim ent Station. Past-President of the
A.S.M.E.
4 Numbers in parentheses refer to the Bibliography a t the end of
the paper.
Contributed by the Special Research Committee on Critical Pressure
Steam Boilers and presented at the Annual Meeting, December
4 -8 , 1939, Philadelphia, Pa., of T h e A m e r ic a n S o c i e t y o f
M e c h a n i c a l E n g i n e e r s .
N o t e : Statem ents and opinions advanced in papers are to be
understood as individual expressions of their authors, and not those
of the Society.
not considered.
677
In 1936, K. Sigwart (7) published data on the viscosity of water
and steam as determined by a capillary tube. The results for
the viscosity of water are in close agreement with those reported
by the authors. However, the values for the viscosity of superheated
steam differ widely from the results reported by the
authors (3) and, in some cases, disagree by as much as 3 to 1.
A comparison of the results of several different determinations
of the viscosity of steam is shown in Fig. 1. H. Speyerer (8) has
recently published a discussion of the conflicting results which
have been reported by the various investigators. The wide
divergence in results has led to a careful examination of the work
of Sigwart (7) as well as of the apparatus, the teohnique, and the
results of the authors; also new determinations of the viscosity
of superheated steam have been made and are reported in this
paper.
Essentially, the method used by the authors in their earlier
investigation was to measure the time of fall of a cylindrical body
in an accurately bored cylinder containing a fluid of known viscosity
and to measure the time of fall of the same body in the
same cylinder containing water or steam under a known pressure
and temperature. It is necessary that the flow of the fluid in the
Temperature, F
F i g . 1 V is c o s it y o f S t e a m a s R e p o r t e d b y R e c e n t I n v e s t ig a t o r s
annular space between the falling body and the cylinder be in the
viscous region. This flow was limited in the experiments to a
maximum Reynolds number of 2100 based on the calibration
data. The essential data required for a viscosity determination
by this method are the density of the fluid under test, the density
and viscosity of the calibration fluid, and the time of fall. I t is
not difficult to measure pressure and temperature and thus compute
density with reasonable accuracy. The viscosity of the calibration
fluids was taken from the International Critical Tables.
The time of fall was recorded automatically on the strip chart
of an Esterline-Angus recording milhammeter in a manner which
eliminated the human element. It is believed that any errors introduced
into the results were due not to errors in measuring
pressure, temperature, or time, but to incorrect assumptions
made in developing the theory of the Lawaczeck viscometer or
due to failure to work with pure steam.
In developing the theory of the Lawaczeck viscometer, the
possibility of flow disturbances in the wake of the fall-body was
Recent findings in the field of hydrodynamics

678 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
indicate that such flow disturbances may exist. Considerable
effort has been expended in a theoretical analysis of the problem
of tail turbulence and the magnitude of the error due to the neglect
of this factor. Up to date, a satisfactory solution of the problem
has not been worked out which would permit a separation of the
resistance to falling into viscous drag in the annular space around
the fall-body and turbulence at the tail of the fall-body. Also,
it may be that, under certain conditions, the motion of the fallbody
is not translatory but vibratory.
Sigwart (7) used the capillary-tube method of measuring the
viscosity of steam and water. Two capillary tubes were used,
each about 14 in. long. The quartz capillary had an internal
diameter of 0.375 mm, while the internal diameter of the platinum
capillary was 0.547 mm. Steam was generated in a small boiler,
superheated, passed through the capillary, throttled, condensed,
and weighed. Sigwart states that a brown precipitate formed in
the condensate after exposure to the atmosphere for some time.
The differential pressure across the capillary was measured by
means of a differential-pressure ring balance (9). Sigwart reports
only one short table of original steam data. In this table,
the maximum and minimum heads, due to the flow of steam in
the capillary, were 0.247 and 0.184 in. Hg, respectively. It is
obvious that the accuracy and sensitivity of the balance must
be of a very high order if dependable results are to be obtained
from such small heads.
The differential head, measured by the ring balance, was reported
by Sigwart as being due to the equivalent of (a) flow
through the capillary of the fluid leaving the system, and (b)
flow of fluid from one leg of the manometer through the capillary
and connecting tubing to the other leg of the manometer.
The connections from the manometer to the capillary
were mounted in a horizontal plane. It is obvious that, when
measuring the viscosity of superheated steam, these manometer
legs contained superheated steam of varying temperature, density,
and viscosity; saturated steam and water of varying temperature,
density, and viscosity. Based on certain assumptions, which did
not include the variable physical state of the fluid in the manometer
legs or the length of tubing occupied by the fluid in these
various states, Sigwart derived an equation from which he estimated
the time elapsed from the opening of the throttle valve on
the discharge side of the capillary until the pressure drop, as
measured by his balance, equaled at least 99 per cent of the head
loss in the capillary. He then took the manometer reading at
this time interval, computed a correction, and used this value to
calculate the viscosity. Thus his readings were taken before a
steady flow state had been established.
R e q u ir e m e n t s o f a C a p il l a r y f o r M e a s u r i n g t h e V is c o s it y
o f S u p e r h e a t e d S t e a m
In order to check the results of their earlier investigation by
a different method and to compare the new results with those
of other investigators, the authors decided to use a capillary
tube designed and operated in accordance with the following
principles:
1 The capillary tube must have sufficient length to give a
head which could be measured accurately.
2 The length of the capillary should be great enough so that
the kinetic-energy correction at the ends would be negligible.
3 The tube should be made of noncorrosive material of
sufficient strength to prevent distortion at the highest operating
pressures and temperatures.
4 A uniform temperature must be maintained along the entire
length of the capillary.
5 The manometer must be connected to the capillary in such
a way that separation between the liquid and vapor phase of the
fluid would be in the horizontal plane of the capillary.
6 In order to eliminate the human element, the manometer
must be designed and operated to give a permanent photographic
record of the head.
7 No readings should be taken until a steady flow state had
been attained and the record of manometer readings must indicate
the presence of a steady flow state for a test period of at
least 1 hr.
8 The theoretical constant of the capillary should be checked
at periodic intervals during the investigation and at the conclusion
of the tests by the use of suitable fluids of known viscosity.
D e s c r i p t i o n o f A p p a r a t u s
In accordance with the above requirements, the capillary was
made from 103.2 lin ft of 0.25-in. outside diam X 0.09294-in. inside
diam seamless nickel tubing. The random lengths of tubing
were connected into one continuous length of uniform internal
diameter by facing each end square, butting it tightly against the
end of the next piece in a snugly fitting nickel sleeve and welding
the sleeve to the tube ends. Each joint was radiographed after
welding to make sure that the ends of the tubes were in alignment
and fitted against each other to form a smooth internal surface.
The capillary was wound into a coil having a diameter of 2.5 ft,
covered with a uniformly spaced electric-resistance-heater wire
and thoroughly insulated. During the course of the investigation,
the capillary was also wound into circle bends with a 1-ft
radius, connected by straight runs about 12 ft long, and was later
coiled on a 2-ft diam.
The manometer, made of two vertical legs of extra-heavy
'A-in. iron-pipe-size, seamless steel tubing, was connected to the
junction blocks at the ends of the capillary by tubing which was
mounted in the horizontal plane of the ends of the capillary, in
order to avoid any errors due to the variation in the level of the
plane of separation of water and steam between the capillary and
the manometer. The manometer was filled half full of mercury
with water above the mercury level. Care had to be used to
eliminate all air from the manometer legs and connections during
the filling process. An X-ray tube was used to radiograph the
mercury levels in the manometer and leave a permanent record
on the film which was mounted in back of the tube as shown in
Fig. 2. Heads up to 5.5 ft could be measured with this apparatus.
The boiler was supplied with deaerated distilled water by placing
two tanks, A and B in Fig. 2, between the feed pump and the
boiler. The tanks were connected by pipe at their upper ends

HAWKINS, SOLBERG, POTTER—THE VISCOSITY OF SUPERHEATED STEAM 679
and were filled half full of xylol. This liquid will not mix with
water nor contaminate it in any way at low temperatures. Before
starting a test, boiling distilled water was supplied to the
lower end of tank B and most of the xylol was transferred to tank
A. Then during the test, the pressure from the boiler-feed pump
was transferred to the feedwater through the xylol. In this way,
only pure distilled water entered the boiler. During many of
the tests, the xylol tanks were by-passed and the boiler was fed
directly from the pump. The results were not affected by this
change in the feedwater system.
The boiler consisted of a short length of double-extra-heavy 2-
in. seamless steel pipe closed at each end by welding, set on an
incline, and fired by a battery of Bunsen burners. Steam was
superheated by placing an electric heater around a section of
nickel tubing. The input to this heater was controlled automatically
by a Wheelco Capacitrol and magnetic switch in order
to maintain a constant steam temperature at the entrance to the
capillary. Two calibrated thermocouples, welded into the bottom
of wells, were placed opposite each other in a thermocouple block
immediately ahead of the capillary (7\ and 7’2 in Fig. 2) in order
to determine the steam temperature entering the capillary.
Twenty thermocouples were provided for reading temperatures
along the capillary, and the heater winding was adjusted to maintain
a uniform temperature from end to end. The steam leaving
the capillary was throttled, condensed in a copper coil, and
weighed.
T h e V is c o s it y E q u a t io n f o r t h e C a p il l a r y
Poiseuille’s law for viscous flow in a straight capillary, with
allowance for kinetic-energy losses at entrance and exit, may be
stated as follows
where n = viscosity
r = radius of capillary
A P = pressure difference between the ends of the capillary
Q = volume of flow per unit of time
I = length of capillary
M = a constant
p = density of the fluid
X = a constant
For long straight capillary tubes in which the end kinetic-energy
corrections may be neglected, the equation is reduced to the form
I = 103.2 ft
D = 0.09294 in.
w = weight flow rate in lb per sec
v = average specific volume in ft3 per lb
ix = viscosity in centipoises
Then C = 26.7276 X 10~» or
The capillary, 103.2 ft long, was first coiled on a 2.5-ft diam.
Because of serious heater trouble, this arrangement was abandoned
and the capillary was formed into circle bends having a 2-
ft diam, connected by straight sections 12 ft long. Of the total
length, 69.6 per cent was straight and 30.4 per cent curved to a 1-
ft radius. This capillary arrangement has been called coil I in
this report. Continued heater trouble made it necessary to obtain
a new heater and to coil the entire capillary on a 2-ft diam,
called coil II in this report.
For a bent capillary, the work of White (10) and Keulegan and
Beij (11) indicates that at Reynolds’ number above approximately
200, the effect of curvature increases the resistance to flow
and the capillary constant then becomes a function of the Reynolds
number and the ratio of the diameter of the capillary to the
diameter of the coil. Equation [3] for the straight capillary may
be modified to account for curvature by introducing a factor

680 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
where C

HAWKINS, SOLBERG, POTTER—TH E VISCOSITY OF SUPERHEATED STEAM 681
At pressures above 1500 lb per sq in. abs, the high pressure
drop through the throttle valve and the low specific volume of the
steam made it virtually impossible to maintain steady flow conditions
for a minimum test period of 1 hr. Inasmuch as the
maintenance of a steady flow state for at least 1 hr, as evidenced
by successive radiographic exposures of the manometer, was
specified in advance as one of the criteria to be fulfilled before any
test data would be accepted, no tests are reported in this paper
for pressures above 1810 lb per sq in. abs.
In 1933, R. Plank (12) presented theoretical equations for the
viscosity of gases and vapors. In the development, he showed
that the thermodynamic equation of state may be expressed

682 TRANSACTIONS OF TH E A.S.M.E. NOVEMBER, 1940
F ig . 7 T h e V is c o s it y o f S t e a m
(1939 P urdue U niversity values.)
Equation [16] gives results which are in agreement with the
smoothed values reported in this paper as shown in Table 5.
Except for the 100-lb per sq in. abs pressures, the maximum
deviation between the computed and smoothed results is 3 per
cent. For the case of 100-lb per sq in. pressure the deviation is
less than 3 per cent for temperatures ranging from 700 F to 1000
F but increases to 6.5 per cent at 400 F. The constants of Equation
[16] are based upon the smoothed values reported in this
paper for a maximum pressure of 1500 lb per sq in. abs, and it is
probable that, at higher pressures, the variation in the specific
volume of superheated steam is too complex to be expressed by
the three terms of the equation.
Values of absolute viscosity have been obtained from Fig. 7
and tabulated in Table 6, together with the computed kinematic
viscosity and several convenient conversion factors.
The viscosity data reported in this investigation are considerably
lower at the high pressures than those presented by the
authors in 1935. The higher 1935 values are probably due to end
effects and wake turbulence of the fall body in the Lawaczeck
viscometer.
A c k n o w l e d g m e n t s
The authors are indebted to Mr. G. H. Van Hengel of The Detroit
Edison Company for constructive criticisms and valuable
assistance rendered in analyzing and interpreting the data of the
investigation, particularly with reference to the effect of curvature

HAWKINS, SOLBERG, POTTER—TH E VISCOSITY OF SUPERHEATED STEAM 683
of the capillary. Dr. Max Jakob of the Armour Institute of
Technology has been most helpful in connection with the solution
of certain problems encountered in analyzing the results.
B IBLIOGRAPHY
1 “ Die Bestimmung der Zahigkeit des W asserdampfes,” by H.
Speyerer, Forschung auf dem Gebiete des Ingenieurwesens, no. 273,
V.D.I., Berlin, p. 1925.
2 “ Bestimmung der Zahigkeit von W asserdampf,” by W.
Schiller, Forschung auf dem Gebiete des Ingenieurwesens, vol. 5, March-
April, 1934, pp. 71-74.
3 “ The Viscosity of W ater and Superheated Steam ,” by G. A.
Hawkins, H . L. Solberg, and A. A. Potter, Trans. A.S.M .E., vol. 57,
Oct., 1935, pp. 395-400.
4 “ Uber Zahigkeit und Zahigkeitsmessung,” by F. Lawaczeck,
Zeit. V.D .I., vol. 63, 1919, pp. 677-682.
5 “ Eksperim ental’nie Dannie i Interpolyatsionnya Formula Dlya
Vyazkosti Vodi,” by V. Schougayew, Vsesoyuzny Teplotekhnicheskiy,
Institut, Izvestiya, 1934, no. 7, pp. 47—49.
6 Discussion of paper “ The Viscosity of W ater and Superheated
Steam ,” by G. A. Hawkins, H . L. Solberg, and A. A. Potter, Trans.
A.S.M.E., vol. 58, April, 1936, pp. 258-262.
7 “ Messungen der Zahigkeit von Wasser und W asserdampf
bis ins kritische Gebiet,” by K. Sigwart, Forschung auf dem Gebiete
des Ingenieurwesens, vol. 7, M ay-June, 1936, pp. 125-140.
8 “ Zahigkeit von W asserdampf,” by H. Speyerer, Forschung
auf dem Gebiete des Ingenieurwesens, vol. 10, March-April, 1939, pp.
100- 102.
9 “ Messung kleiner Druckunterschiede bei hohen absoluten
Drucken,” by E. Schmidt, Zeit. V.D .I., vol. 80, 1936, pp. 635-636.
10 “ Streamline Flow Through Curved Pipes,” by C. M. White,
Proceedings of the Royal Society of London, series A, vol. 123, 1929,
p. 645.
11 “ Pressure Loss for Fluid Flow in Curved Pipes,” by G. H.
Keulegan and K . H. Beij, National Bureau of Standards, Journal of
Research,_vol. 18, Jan., 1937, p. 965.
12 “ Uber die Zahigkeit von Gasen und Dam pfen,” by R. Plank,
Forschung auf dem Gebiete des Ingenieurwesens, vol. 4, January-
February, 1933, pp. 1-7.
D iscussion
E. F. L e i b .6 For the analytical representation of their data
on the viscosity of superheated steam, the authors rely on Equation
[14], as recommended by R. Plank. This equation has been
derived from the kinetic theory of gases for the viscosity at high
densities, where the space occupied by the molecules is of the
order of magnitude of the gas volume, and the transfer of momentum
occurs at collisions as well as between collisions. Definite
values have been derived for all constants in Equation [14] and,
due to their significance, these constants are essentially positive.®
However, the constants in the authors’ Equation [16] have
nothing in common with these theoretical values and one constant
is even negative. Therefore, the latter equation becomes
purely empirical. The reason for the failure of Equation [14] is
that, in the range investigated by the authors, the state of the
vapor is that of moderate density, where the volume of the molecules
is of minor importance, and the increase of the viscosity
over the corresponding gas condition is mainly due to the forces
of attraction between the molecules.
The viscosity of a vapor can, in first approximation, be obtained
as follows: From the kinetic theory, we have the fundamental
relation for the viscosity
whence the viscosity of the perfect gas, with I = la is
mo = Vspco lo........................ ............................ [19]
where Co is the mean velocity of the molecules due to the thermal
agitation.
The kinetic theory gives
where g = acceleration of gravity, R = gas constant, D = diameter
of molecule, b = van der Waals’ sphere of exclusion, v =
specific volume. Thus, mo depends only upon the temperature T.
Assume the density to be such that only binary encounters
between the molecules occur. Consider a molecule midway between
two collisions, where it will travel with the velocity Co.
While approaching the object of its next collision, it enters the
field of attraction and experiences an acceleration. Thus, the
velocity will be greater than c0, and then, according to Equations
[18] and [17] of this discussion, the number of collisions and
hence the viscosity will be greater than in the gas condition.
To calculate this acceleration effect, we use a method similar
to that by which Sutherland investigated the effect of the attraction
on the mean free path. If u(r) is the potential energy of a
molecule with respect to another molecule in the distance r, the
force of attraction is
Then the motion of a molecule subjected to this force is given by
6 Combustion Engineering Company, Inc., New York, N. Y.
6 “ M athem atical Theory of Non-Uniform Gases,” by S. Chapman
and T. G. Cowling, Cambridge University Press, London, 1939, p. 288.

684 TRANSACTIONS OF TH E A.S.M.E. NOVEMBER, 1940
The kinetic theory gives for the mean free path at moderate
densities
Equation [34] is the relation between the viscosity at moderate
densities and the viscosity in the gas condition at the same temperature.
The quantity 6 can be identified approximately with
the smallest volume which the liquid substance can assume without
being compressed; i.e., for H20 , b = ~0.0160 cu ft per lb.
The quantity 0 was taken as an average calculated from Equation
[34] with the 1939 Purdue test data of the viscosity of steam
and was obtained as
9 = 19,000 F
With these constants, Equation [34] represents the test data
with a deviation of a few per cent only. Due to the derivation
from kinetic conceptions, this equation can be expected to represent
the viscosity of steam far beyond the 1500-lb per sq in. limit
reached in the 1939 Purdue experiments.
Equation [30] permits an exact integration for n — 2. Though
the actual value of n may be greater, the error introduced by the
choice n = 2 will be noticeable only at high densities, since it
corresponds to a less rapid decrease of the potential energy with
the distance between the molecules. Any value of n greater than
unity is compatible with the equation of state for vapors, as derived
by the writer from statistical methods, due to the particular
type of molecular clusters chosen.7
The effect of this choice on the viscosity at moderate densities
is negligible, and we obtain by integration of Equation [30], between
the limits described, for n = 2
7 “Thermodynamic Properties of Vapors,” by E. F. Leib, presented
at the Spring Meeting, Worcester, Mass., May 1-3, 1940, of The
American Society op Mechanical Engineers.
G. H . V a n H e n g e l .* Recently, Russian results9 have been
published which differ with the results of the authors and other
previous investigators. The many differing test results by all the
various experimenters indicate the great difficulty in determining
the viscosity of steam (8).10 I t seems that the capillarytube
method has inherent difficulties, some of which the writer
will here review.
First let us consider the mechanical difficulties of the capillary.
The authors have eliminated the complication of the entrance and
exit losses of the capillary by taking a long capillary. This did
away with the corrections investigated by Schiller.11
They introduced, however, some other problems, viz., the
effect of the curvature of the coil and the problem of determining
the tube diameter. Accurate determination of the diameter is a
A pD5
very important one as X = 25.103 ------- for a length of 103.2 ft
W2 v
8 Mechanical Engineer, The D etroit Edison Company, Detroit,
Mich. Mem. A.S.M.E.
9 “Viscosity and H eat Conductivity of Steam for High Temperatures
and Pressures,” by D. L. Tim rot and N . B. Vargaftik (in
Russian), Journal of Technical Physics U.S.S.R., vol. 9, no. 6, 1939,
pp. 461-469.
10 Numbers in parentheses refer to the Bibliography at the end of
the authors’ paper, page 683.
11 “ Untersuchungen iiber Laminare und Turbulente Stromung,"
by L. Schiller, Forschungsarbeiten a u f dem Gebiete des Ingenieurwesens,
heft 248, 1922.

HAWKINS, SOLBERG, POTTER—THE VISCOSITY OF SUPERHEATED STEAM 686
and Re is inversely proportional to D, where X = a dimensionless
resistance coefficient, D = inside diameter in ft, W = weight flow
rate in lb per sec, and v = specific volume in cu ft per lb.
The problem of the coil curvature has been investigated (10)
and (11) but none had a coil of such a length or of such a heavy
wall. Becauge it forms an oval cross section, the bending of a
thick-wall tube may affect the tube diameter and cross-sectional
area far more than previous investigators found.12 This deformation
reaches larger proportions at higher internal pressures, an
indication of which can be found in an article by Dean.13 This
deformation may even affect the X vs. Re curve for the coil.14
The determination of the inside diameter of the tube presented
another problem with such a long coil. The inside diameter
actually measured before the bending was 0.091 in. but the inside
diameter after the bending could not be measured directly. The
intemal-volume determination with a liquid gave too large a
diameter, due to the filling up of the connections between the different
tube lengths. Gas-volume determination did not seem to
offer higher accuracy.16 Thus, finally, the tube diameter of
0.09294 in. was established by using White’s formula1* and the
authors’ 1940 calibration tests with water shown in Table 2 of the
paper under the assumption of a round capillary. Deviations up
to 5 per cent occurred for unknown reasons in the X values.
In the paper, the influence of the temperature on the expansion
of the nickel tube has been neglected, although this affects the
computation of the viscosity by 3 per cent at the highest temperatures
of 1000 F.
Another difficulty is brought out by Keulegan and Beij17 that
a little sag in a straight capillary contributes to large changes in
friction coefficients. Here the coil was wound at random and was
not at all the case of a coil in a horizontal plane.
The effect of the foregoing deviations, on which the viscosity
calculation of steam is so vitally dependent, should have been
established with the calibration test before the final steam tests
could proceed.
But these mechanical deviations may not be the only items of
great importance, because of other difficulties presented by thermodynamic
problems. In the first place, among these is the
effect of nonisothermal conditions in the viscous flow of incompressible
fluids on the friction coefficient. ,,>1* The slightest heat
transmission will affect the parabolic velocity pattern considerably
and thus influence the friction coefficient.
In the case of compressible fluids, there is added to the foregoing
problem yet another one, i.e., the knowledge of the proper
relationship between the pressure drop and the friction coefficient.
This relationship will be different for the isothermal and
adiabatic cases.20
12 Bibliography (11), p. 96, and (10), p. 652.
13 “ D istortion of a Curved Tube Due to Internal Pressure,” by
W. R. Dean, Philosophical Magazine and Journal of Science, vol. 28,
Oct., 1939, pp. 452—464.
14 Bibliography (10), Fig. 6, p. 657.
15 “ Determ ination of the Internal Volume of Steel Capillaries for
Measurements W ith Gases,” by J. Kam insky and B. E. Blaisdell,
Review of Scientific Instruments, vol. 10, Feb., 1939, pp. 57-58; see
also correction, ibid., May, 1939, p. 151.
16 Bibliography (10), p. 663.
17 Bibliography (11), final paragraph, p. 98.
18 “ Pressure Drop and Velocity Distribution for Incompressible
Viscous Non-Isothermal Flow in the Steady State Through a Pipe,”
by A. Lee, W. D. Nelson, V. H. Cherry, and L. M. K. Boelter, Proc.
Fifth International Congress of Applied Mechanics, John Wiley &
Sons, Inc., New York, N. Y., 1939, pp. 571-577.
'* Uber den W armeaustauch bei der Stromung zither Flussigkeiten
in Rohren,” by M. Jakob and H . Eck, Forschung auf dem Gebiete des
Ingenieurwesens, vol. 3, no. 3, 1932, p. 121.
20 “ Steam and Gas Turbines,” by A. Stodola and L. C. Loewenstein,
vol. 2, McGraw-Hill Book Company, Inc., New York, N. Y.,
1927, p. 1026.
F i g . 8 C o m p a r is o n o f P u b d u e T e s t V a l u e s W i t h t h e 1939
P u r d u e S m o o th e d V a l u e s o f t h e A b s o l u t e V is c o s ity o f S te a m
Perhaps one or more of these questions may play an important
role in interpreting the results because, if the data of
the authors’ Table 4 were plotted in Fig. 7 of the paper, as in the
case in the writer’s curve Fig. 8, it will be noted that several
points have a viscosity which deviates more than 10 per cent from
the smoothed values given as curves in Fig. 7. This may be
partly due to the two different coils that were used, but as long
as that effect is not established other causes may exist.
Checking up Plank’s formula results with the final smoothed
values is only a m atter of determining constants and is of no great
importance in view of the deviation of the smoothed values from
the actual test results. The smoothed values should therefore
be used with great reservation.
For the reasons mentioned, the precise method of determining
steam viscosities with capillaries must be more fully examined
and developed before we can consider the values obtained as final.
Now that the authors are equipped with the experience of their
published experiments, let us hope that they will once more go
into this problem of fixing the values of the viscosity of steam.
For pressure-drop calculations, the authors’ values will be useful,
but for calculations of nozzle flow, heat transfer, and the like, a
higher accuracy is desirable.
F. G. K e y e s . 21 Shortly after the publication of the measurements
of steam viscosity by the authors (3) over a wide range of
temperature and pressure (420 C:240 atm), a paper appeared
by Sigwart,22 reporting measurements considerably lower than
those obtained by the American investigators. Recently the
first investigators23 have again measured the viscosity properties
of water using the “capillary-flow” method of Poiseuille, whereas,
the earlier results were obtained by observing the time of fall of a
metal cylinder in a tube filled with steam. Both of these investigations
were carried through with care and with the use of ingenious
controls. The results are in substantial agreement,
albeit the capillary-tube results do not extend over the wide pressure
range reported upon in 1935. It is now established that the
low results of Sigwart do not represent the true viscosity of
steam.
The object of the present comment is to indicate that the entire
ensemble of results by the authors may be simply correlated
as a temperature function of the low-pressure viscosity of steam
plus a function of the pressure only.
There are few data available for the viscosity properties of
21 Research Laboratory of Physical Chemistry, Massachusetts
Institu te of Technology, Cambridge, Mass.
22 “ Messungen der Fahigkeit von W asser und W asserdampf bis
ins kritische Gebiet,” by K. Sigwart, Forschung auf dem Gebiete des
Ingenieurwesens, vol. 7, 1936, pp. 125-140.
23 Preprint of this paper as presented at the 1939 Annual Meeting.

686 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
F i g . 9 S t e a m V isc o s it y R e s u l t s F kom 1935 T e s t s C o m pa r e d W it h T h o s e o f 1939
gases over any considerable pressure range24 and the theory along
classical lines has been mainly developed by Enskog26 to a first
approximation employing the van der Waals concept of the molecular
field. The expression obtained for the viscosity rj is
The constants of Equation [36] were obtained by least-square
procedure using all available low-pressure data.27
Equation [35] may be transformed to the form
rjo is the viscosity as a function of temperatures at low (zero)
pressure and t = (1 + 5/8 b/v + 0.2869 62/r 2); the symbol 6
denoting the van der Waals quantity, 4 times the volume of the
molecules. The formula was applied successfully to the data for
C 0 2.
From the computational point of view, a formula in the variables
T and v is less convenient for engineering purposes than one
in T and p. The writer accordingly attempted in a wholly empirical
way to obtain such a formula and it appears that the
following equation28 represents the authors’ results
V = >70 + ( 0 . 0 3 1 0 3 — 3 . 6 5 X 1 0 - ' p ) p X 1 0 “ * ................[ 3 5 ]
In Equation [35], the viscosity is expressed in poises and the
pressure in atmospheres (14.696 lb-in.-2). The quantity ijo is
the viscosity of steam for pressures approaching zero. The
quantity vo can be represented by the following equation commonly
referred to as the Sutherland formula
, . N 1.851 X 10-‘( D ‘A
vo (poises) = .................... [36]
24 “ The Viscosity of Gases at High Pressures,” by R. O. Gibson
and H. J. Paris, Amsterdam, 1933.
25 "K inetic Theory of H eat Conductivity, Viscosity, and Diffusion
in Certain Condensed Gases and Liquids,” by D. Enskog, Kongl.
Svenska Vetenskaps-Akademiens Handlingar, Stockholm, vol. 63, no.
4, 1922, p. 63.
26 “ Thermodynamic Properties of Steam ,” by J. H. Keenan and
F. G. Keyes, John Wiley & Sons, Inc., New York, N. Y., 1936.
It is clear that if Equation [35] is satisfactory, the function Z
computed from the data should be linear in the pressure. Fig. 9
of this discussion shows the 1935 data which extend to 240 atm
and indicates that the linear relationship is satisfied. There is
moreover no detectable temperature influence left unaccounted
for.
The new data lie in general below the older results but are confirmatory
in general of the older measurements. The upper left
portion of Fig. 9 shows the later data Z function plotted for constant
pressures with temperature as the abscissas. There is no
pronounced trend.
The upper right portion shows the data for low (zero) pressures
plotted as follows: Equation [36] may be written in the form
T1/ ’ 1 c
— = — |— T-1. If the square root of the absolute tempera-
770 a a
ture is divided by the viscosity, the quantity should vary linearly
in reciprocal absolute temperature.28
The experimental data are seen to be adequately represented
by the Sutherland formula of which the constants in Equation
[36] were determined by least-square procedure. The dotted line
represents in the same coordinates the low-temperature viscosi-
27 "A n Experim ental Study of the Viscous Properties of W ater
Vapor,” by C. J. Smith, Proceedings of the Royal Society of London,
Series A, vol. 106, 1924, pp. 83-96.
28 “ The Sutherland Viscosity Constant and Its Relation to the
Molecular Polarization,” by F. G. Keyes, Zeit. fa r Physikalische
Chemie, Cohen-Festband, vol. 130, 1927, pp. 709-714.

HAWKINS, SOLBERG, POTTER—THE VISCOSITY OF SUPERHEATED STEAM 687
ties computed from the formula (10.74 + 0.0178*) employed by
the authors.
S u m m a r y
1 The authors’ new viscosity data do not indicate that any
change in the earlier viscosity Equation [35] is necessary.
2 Equation [35] reproduces the viscosity of steam in relation
to all the authors’ results within the limits of experimental error.
C o m m e n t s
The foregoing discussion was prepared on the basis of the viscosity
of steam data given in Table 6;23 these results extend to
about 123 atm. It is intended to show that the results are in
tolerable agreement with the 1935 data obtained by a method
different from that employed in 1939. The 1935 data extend to a
pressure of 239 atm or nearly twice the range of the later results.
Recently, corrections have been applied to these data,23 resulting
in a considerable reduction in their magnitude. For example
at 801 F and 123.2 atm, 3.83 X 10-4 poises is the corrected
1940 result obtained by correcting the December, 1939, datum,
which is 5.34 X 10“ 4 poises; roughly 40 per cent greater. It is
therefore clear, since the uncorrected 1939 results are in fair
agreement with the 1935 results, that a serious error affects the
1935 data. The authors note this fact.
An examination of the new data using the form of Equation
[35] indicates that the corrected data may be represented within
the limits of accuracy. The formula is as follows
v = i)o + (0.0151 — 5.9 X 10-“ patm) pBtm X 10-4. . . . [38]
The formula for yo is the same, i.e., Equation [36].
If the viscosity of steam at 239 atm 788 F is computed using
Equation [38], the number 2.67 X 10“ 4 is obtained. The 1935
result is however given as 7.30 X 10~4 which would imply that
the Lawaczack method gave a result nearly 190 per cent greater
than that which the capillary method might give. The extrapolation
using Equation [38] to 239 atm (nearly twice the range
upon which this equation is based) may of course be subject to
great error.
normal component velocity is m/3 . Then it is as if a stream of
fluid having the mass velocity mnu/ 2 were flowing squarely
against the unit surface with velocity m/3 and momentum
mnu2/ 6. This stream is brought to zero velocity by impact and
then completely reversed by elastic rebound; an action made
possible by the mechanism of real molecular motion, to which the
stream concept is merely an over-all equivalent. The effect of
impact is now a unit pressure of the value
p = mnu2/3 ....................................[39]
Since mn is the same as density p, the mass per unit of volume and
equal to \/v, the relation becomes
p = pu2/S pv = m2/ 3 ................................[40]
This pv is magnitudinally equal to 2/ 3 of the kinetic energy of all
the molecules in the unit of mass, but physically it is an inertiaforce
effect. The final step of making absolute temperature T
proportional to u 2 goes beyond the needs of the present discussion.
In a gas mixture, molecules of different mass m are put on an equal
basis by the fact that mean energy mu2/ I must be the same for
all molecules at a given temperature.
R. C. H . H ec k.28 The object of this discussion is to set up an
argument in terms of simple kinetic theory, for the necessary increase
of gas viscosity with pressure and with temperature.
As to the detail of its mechanism, viscous action within a gas
is distinctly different from that in a liquid. In the liquid, the
shear stress of viscous resistance and the compressive stress of
pressure reaction are both phenomena of molecular contact;
but, in the gas, they are impact effects of molecules in free translatory
motion. Of molecular motions and interactions in the
liquid phase we have but vague concepts, knowledge (of over-all
kind) coming wholly by observation and empirical determination.
The free and simple motions of gas molecules can be reasoned
in large degree; at least far enough to account for the major
component of behavior that will yet require experiment for close
determination.
First comes a brief outline of a part of the course of reasoning
that leads to the gas equation pv = RT, taking the simplest case
of a single gas with molecules all of the same mass m, having the
mean velocity u. Imagine a cubic unit of gas enclosed by walls
that are perfectly elastic and infinitely fine in structure, so as to
be perfectly flat even to molecules. With n molecules in this
cube and velocity u in terms of the edge of the cube as linear unit
and per second, the number of impacts per second on the square
side is nu/2. These impacts are in all directions, but the mean
21 Research Professor of Mechanical Engineering, Rutgers University,
New Brunswick, N. J. Mem. A.S.M.E.
Considering gas in motion, i.e., in mass motion, three forms of
simple streamline flow are to be distinguished, as follows:
1 Ideal uniform flow, as in a “frictionless” straight channel,
with progressive velocity V the same at all places in the current.
Completely random molecular velocity continues to be distributed
throughout the moving body of gas just as if this were standing
still; and the total or absolute velocity of each molecule (with
reference to the fixed channel) is the vector sum of the particular
m within the current and the general V of the current.
2 Growing velocity of flow, as in current acceleration or jet
formation, with V increasing and mean u decreasing as pressure
falls. Flow velocity V is uniform over surfaces of like state across
the current.
3 Viscous or “laminar” flow, in which V is constant along any
line of flow but varies across the current. This is the kind of
motion in which we are now interested.
In laminar flow, consider a surface of equal velocity V, which
is a cylinder in either of the viscosimeters used at Purdue. This
surface is now to be purely geometrical, offering no impedance to
the passage of molecules. Across the surface, momentum is
being carried at the rate mnu2/6 figured for pressure effect. Also,
across the surface (with dimension y in that direction) there is
the velocity gradient dV/d-y.
Next, consider the cross-stream of molecules from lower to
higher V; these must be accelerated in the flow direction, hence,
exert an inertia reaction against the higher-velocity stream.

688 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
Similarly, the cross-stream from higher to lower velocity must be
retarded, with a forward pulling inertia effect. The two reactions
combine to oppose relative motion of the parallel-flow streams;
and the amount of this opposition, due to a deflection of momentum
on account of gradient dV/dy, is proportional to the total
momentum which manifests itself in pressure p. Since in any
rate of streamline flow, velocity V will be very small relative to
molecular u, viscosity is a very feeble force action in comparison
with gas pressure.
In Fig. 10 of this discussion is shown a chart of absolute viscosity
ii, interpolated from Figs. 9 and 10 of the 1935 paper (3).
The stage of argument thus far reached, with the conclusion
(partial) that n should be proportional to V, would require that
the constant-pressure lines be horizontal and that they be equally
spaced vertically. I t becomes necessary then to perceive some
further influence which will account for the actual behavior seen
in Fig. 10 of this discussion; and this is found in the matter of
mean length of path of molecule between impacts. The modifying
effect of path length is argued thus:
With rise of pressure along the isotherm, the increase of density
p shortens the mean path; with a shorter path across the flow
direction or in the direction of velocity gradient, there is smaller
change of the V component of the total molecule velocity; and
therefore the increment of m with p lessens as the pressure is
higher.
Along a p-constant line, smaller density with higher temperature
lengthens the path of the molecule, hence there is greater
change of momentum in the direction V, and n rises with the
temperature.
In Fig. 10 the dot-and-dash curve is the handbook curve of m
for air, at or near atmospheric pressure.
From the simple reasoning thus completed it is not possible to
make a quantitative prediction; it is explanatory rather than
determinative; and a theory of full quantitative capability
would be far more elaborately mathematical. Also, for steam as a
superheated vapor or a gas near to saturation, there is the further
influence of departure from ideal-gas behavior.
[An additional discussion of the paper, as presented in preprint
form at the 1939 Annual Meeting of the Society, was received
from J. R. Finniecome of Manchester, England. Since revisions
were made in the original paper, it has been impossible to communicate
these to him in order that he might have the opportunity
to revise his discussion. Therefore, it is with regret that it
has been found necessary to omit his contribution at this time.
E d i t o r .]
A u t h o r s ’ C l o s u r e
The discussions of Messrs. Lieb and Keyes are very interesting
because different types of equations are developed which closely
represent the smoothed values presented in the paper. Dr. Lieb’s
equation, based on the kinetic theory, should permit extrapolation
of the results to pressures much higher than the upper limit
of the experimental data reported by the authors.
Mr. Van Hengel has contributed a very thorough and valuable
discussion of the difficulties encountered in measuring viscosity
by means of a capillary tube. His remarks should be of great assistance
to any person interested in experimental work in this
field.
Possible deformation of the cross-sectional area of the capillary
on bending to a one-foot radius is one of the mechanical difficulties
mentioned by Mr. Van Hengel. Five sections have been cut
from the capillary coil, the ends have been polished perpendicular
to the coil tangent, and enlarged photographs have been made of
the tube cross section at a magnification of 100 diameters. The
photographs do not show evidence of deformation of the circular
cross section at any of the five sections investigated.
Mr. Van Hengel discusses the difficulty of accurately measuring
the internal diameter of the capillary and points out that the
fifth power of the diameter appears in the general flow equation
of Blasius. The constants in the authors’ Equations [4], [5],
and [12] and the calibration data of Fig. 5 are based on a particular
numerical value of the diameter. If the actual diameter of
the capillary remains constant and the same value for the diameter
is used in all computations involving both calibration and
test runs, the final values for the viscosity of steam as computed
by the method of the authors are independent of the numerical
value used for the diameter.
Mr. Van Hengel mentions the effect of temperature on the diameter
of the capillary. This has been considered in computing
the Reynolds numbers used in connection with Fig. 5, but its
effect upon the curves in Fig. 5 has been neglected. Thus a partial
correction has been made and errors due to neglect of expansion
of the capillary at high temperatures are considerably less
than those indicated by Mr. Van Hengel in his discussion. The
use of the general Blasius equation with friction coefficients in the
form shown in Fig. 6 would have made the correction for expansion
simple and complete but the final results would have depended
upon the fifth power of the diameter of the capillary. To
summarize, then, the method of computation used by the authors
gives viscosity data which are independent of the numerical
value assigned to the diameter of the capillary if this diameter is
constant, but complete corrections for changes in the actual diameter
with temperature, while small, are nevertheless almost impossible
to make. The method preferred by Mr. Van Hengel
of using the Blasius equation makes it possible to correct completely
for changes in capillary diameter with temperature, but
the numerical value of the diameter to the fifth power enters directly
into the results at all temperatures and therefore necessitates
a precise determination of the diameter.
In conclusion, the authors wish to thank all of those who have
presented oral or written discussions for their interest, comments,
and constructive criticisms.

S team -T urbine Blading
This paper reviews the blading-design practice associated
with modern high-pressure high-tem perature steam
turbines. The design problems encountered in the development
of partial-adm ission im pulse blading for topping
units are described, as well as the current engineering
practice employed in the m anufacture of such blading.
The stress analysis used in the construction of fulladm
ission blading is reviewed. The design procedure
adopted for high-tip-speed last-row blading and the
natural lim its in capacity imposed on 3600-rpm turbine
construction are also discussed. Materials for turbine
blading are considered, as well as the m etallurgical problems
associated with the fabrication and welding o f the
high-grade alloy steels now available.
I n t r o d u c t io n
NOZZLES and blades for large steam turbines may be classified
inclusively as elements incorporating a series of curved
passages which control and direct the power-producing
phases of the steam flow through the turbine. The general function
of the steam passages is twofold: (1) The kinetic energy in
the carry-over velocity from the preceding stage must be efficiently
turned through a predetermined angle; (2) at the same
time, there must be an efficiently carried out expansion through
each element by means of which the kinetic energy of the working
fluid is increased.
The foregoing general statement may be taken to cover moving
or stationary elements or blading of the impulse or reaction type.
Reaction-turbine designers often find it desirable to divide the
thermal drops unequally between the moving and stationary
blades of a stage. Impulse-turbine builders find that the highest
efficiency can be achieved by the adoption of a certain amount of
pressure drop in the moving blades. The original sharp distinction
between reaction and impulse turbines has, therefore,
largely disappeared, although there are still notable differences
between the two types with respect to mechanical details. The
present problem of the blading engineer is rather that of deciding
what type of element makes possible the most reliable construction
which also will meet the required efficiency level at the lowest
manufacturing cost.
The research and engineering aspects of the thermodynamic
and fluid-flow problems related to blading design are of the utmost
importance. Extensive research has been carried out and
is still under way in connection with the determination of the
most efficient forms of steam passages. In recent years, attention
has been given to the further development of steam-flow
problems along lines adopted by aerodynamic engineers in studying
the lift and drag of airfoil sections. It is not the purpose of
this paper to present the steam-flow phases of the work, but rather
to indicate some of the mechanical aspects of blading construction
which have been developed in the last few years to meet the
B y R. C. ALLEN,1 MILWAUKEE, WIS.
1 Engineer, Steam Turbine Departm ent, Allis-Chalmers M anufacturing
Company.
C ontributed by the Power Division and presented at the Semi-
Annual Meeting, Milwaukee, Wis., June 17-20, 1940, of T h e A m e r i ­
c a n S o c i e t y o f M e c h a n i c a l E n g i n e e r s .
N o t e : Statem ents and opinions advanced in papers are to be
understood as individual expressions of their authors, and not those
of the Society.
demands for high-capacity high-speed machines, designed for
maximum pressures and temperatures. Some of the considerations
given may seem elementary; others may be open to argument
and discussion. In spite of the fact that the steam turbine
is not a newcomer in the field of engineering, it must be admitted
that recent developments have caused quite radical changes in
design reasoning. It may further be said that the design development
of machines for the highest pressures and temperatures is
still in the process of revision and research.
In this paper, the author has been permitted to supply engineering
data from the steam-turbine practice of the Allis-Chalmers
Manufacturing Company. The subject m atter concerns
principally the recent mechanical developments in blading elements.
BLA D IN G M ATERIALS
A i r - H a r d e n i n g V e r s u s N o n - A i r - H a r d e n i n g S t a i n l e s s S t e e l
Even the advanced metallurgical practice of the present day,
has not produced a blading material in which the most desirable
values of all the various physical and chemical properties can be
provided in the same alloy. For example, in some materials in
which there is high strength at high temperature, the hardness is
so great that they cannot be machined. Other materials may
have excellent properties, except that the corrosion resistance
may not be sufficiently great. Among the materials available,
there will be found certain classes of alloys which do not have the
highest resistance to stress corrosion, but which do have other
essential qualities, i.e., high physical properties at elevated
temperatures and weldability without air-hardening. Consideration
of all these properties indicates that such materials best
meet the requirements for most blading service. In particular,
these statements refer to certain of the alloy steels, in which the
iron constituent is in the gamma phase, which come under the
general classification of austenitic steels.
These materials must be properly processed to avoid undue
carbide precipitation or growth of the crystalline grains at the
steel mill or during the final-assembly heating operations. Under
certain conditions of application, suitably low stresses may of
necessity be adopted if impurities are present in the steam which
might, in the presence of high stresses, cause stress corrosion. It
is the conclusion that the strength at high temperature, the resistance
to the various types of corrosion, and appropriate consideration
of welding make alloys of this class the most generally
useful in steam-turbine-blading practice.
C y c l o p s No. 17-A A l l o y
The standard Allis-Chalmers blading steel is known as Cyclops
No. 17-A alloy and is produced by the Universal Cyclops Steel
Corporation, Titusville, Pa.
The composition is as follows:
-—Per cent—»
C arbon....................................................... 0.35 to 0.45
Nickel......................................................... 19 to 20
Chrom ium ................................................. 7 to 8
Silicon......................................................... 0.95 to 1.25
Manganese................................................ 0.55 to 0.75
This material has an ultimate strength of 105,000 psi at room
temperature in rolled bars 3/< X lVs in. The corresponding
689

690 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
proof stress at 0.01 per cent plastic yield will be about 50,000
psi. The material has good physical properties at high temperatures.
Recent developments with this material have been in the direction
of improved quality and closer control of process work. The
austenitic structure of this alloy and its consequent capability
of being welded without air-hardening facilitates the bladeassembly
processes. During the last year, there has been an
extensive study of the rolling-mill processes employed in the
production of sections which are rolled to final size. In previous
practice, the rolling-mill passes following the last rolling heat
were carried out in such a manner that the percentage reduction
at various parts of the blade section was not the same. The percentage
reduction in the region of the outlet edge was much
greater in the last roll passes than the corresponding reduction in
the thick part of the blade section; hence, there was a substantial
difference in grain shape and size between the material
in the edges and the material in the thick parts of the sections.
Harold Stein2 has developed and made effective during the
last year a change in the rolling-mill program, whereby the
rate of reduction in the passes following the last rolling heat
is made essentially the same throughout the widths of the
blade sections. This procedure is effective in producing grains
substantially equiaxed in form and practically equal in size over
the full widths of the sections. The internal stresses set up between
the thin and thick portions of the blade sections are reduced
considerably by this procedure. It is the present conclusion
that certain cases of stress corrosion will be prevented by
this change in the method of rolling.
In the rolling of blade sections by the older method, the ends
of the bars came out of the rolls tapered to a considerable degree,
the bar being several inches longer on the outlet-edge side than on
the thicker inlet-edge side. With the new process, the material
in the inlet and outlet edges is of approximately the same length.
In other words, the bars come out of the rolls nearly square at
the ends.
Further improvements have been made at the steel mill with
respect to finish-rolling. A cold-finishing pass has been adopted,
which gives a higher degree of surface finish to the sections.
After the final cold pass, the Cyclops material is stress-relieved
at 1425 F.
1 3 -C h b o m e S t a in l e s s I r o n
The analysis of 13-chrome stainless iron is as follows:
Per cent
Carbon................................................. 0.06 to 0.13
Chromium............................................ 11.5 to 13.0
Silicon.................................................. 0.50 Max
Manganese........................................... 0.25 to 0.80
Nickel................................................... 0.50 Max
For the extreme low-pressure end where the operating temperature
is low and where the maximum strength is necessary, the
13-chrome stainless iron is considered the best material available.
It possesses reasonable corrosion resistance, combined with high
strength and good endurance limit. Welding operations must
be carefully planned and executed, otherwise the air-hardening
characteristics of this material may introduce difficulties. The
13-chrome stainless iron is used by all of the large turbine builders
in this country for the highest-stress low-pressure blades. In
Allis-Chalmers practice, this material is used for low-pressure
blades in which the annular-blade outlet area at 3600 rpm exceeds
approximately 21 sq ft.
2 Director of Research, Chemistry, and Metallurgy, Allis-Chalmers
M anufacturing Company, Milwaukee, Wis.
19 P e r C e n t C h r o m e , 9 P e r C e n t N ic k e l S h r o u d M a t e r ia l
This is a standard alloy used for reaction-blade-shroud sections
of the present standard type which are secured in place by welding.
All new central-station machines will have reaction blading
so shrouded.
The carbon content is about 0.1 per cent. The columbium
content varies from 0.7 to 1 per cent to stabilize the material
against carbide precipitation. This material is capable of being
easily welded. Standard embrittlement tests show that the
material is substantially immune to this type of attack.
25 P e r C e n t C h r o m e , 12 P e r C e n t N ic k e l W e l d R od
Coated weld rods of this material containing about 0.1 carbon
with a stabilizing addition of columbium are made up in sizes as
small as V32 in. in diameter. They are used for welding 19-9
chrome-nickel shrouding on the Cyclops 17-A blading, or the welding
together of the integral shrouding on the Cyclops impulse
blades.
The welding of austenitic blading materials requires special
welding units of the mercury-arc-rectifier type which permit welding
with current flows down to about 2 amp. Such welds are
formed with a minimum of local heating. Excessive grain growth
is prevented by the rapid cooling of the weld by the chill effect
of the adjacent material.
E v e r b r it e B a s e M a t e r ia l
Intermediate blading for moderate pressures, temperatures,
and blade speeds is prefabricated in segments in which the ends
of the rolled blade sections are joined by casting them into a
foundation of Everbrite alloy. Everbrite is a casting alloy of the
following composition:
Per cent
Copper..................................................................... 5 9 .0 to 6 5 .0
Nickel............................................................. 2 9 .5 to 3 1 .5
Iro n ................................................................. 5 .0 to 8.0
Manganese.................................................... 0 .6 0 Max
C arbon............................................................ 0 .2 5 Max
Silicon............................................................. 0 .6 0 Max
The process work in connection with this development requires
very accurate control of many factors, including extreme
cleanliness of the blade ends, a carefully worked out process for
the gating and pouring of the blade group castings, and an extremely
close control of the temperature of each pot of metal that
is poured. With proper control, an excellent bond between the
rolled blade material and the cast foundation is obtained.
S u m m a r y o f A p p l ic a t io n o f B l a d in g T y p e s i n S t e a m P a t h
1 Impulse blading is milled from bars of the Cyclops alloy
with integral base and shroud sections. The shrouding is for the
purpose of confining the steam flow to the blade path and also
for reinforcing the structure mechanically. Circumferentialgroove
blading is used up to about 25,000 kw for top-turbine
service and 75,000 kw for condensing machines at 3600 rpm.
Axial-slot roots are used for higher capacities. An important
recent development is the welding of the shrouding in groups
of two blades for large machines.
2 High-pressure high-temperature reaction blading is milled
from Cyclops bars with integral roots. After assembly in the
spindle, the 19-9 chrome-nickel shrouding is welded in place.
A characteristic of the high-temperature shrouding is the small
number of blades per segment, usually not exceeding 4 to 6 for
the highest temperatures.
3 The intermediate blading for moderate temperatures and
stresses is prefabricated in segments by casting the rolled sections
into foundations of Everbrite or cast iron. The 19-9 chromenickel
shrouding is welded in place prior to casting.

ALLEN—STEAM-TURBINE BLADING 691
4 Where stresses make necessary the use of milled low-pressure
blading, the Cyclops alloy is used in conjunction with the
welded 19-9 chrome-nickel shrouding up to annular areas of 21
sq ft at 3600 rpm.
5 Where the annular-blade outlet area exceeds about 21 sq
ft at 3600 rpm, 13-chrome stainless-iron blading is used with
roots of the axial-slot type. Shroud sections and lashing wires
are of 19-9 chrome-nickel steel, welded in place.
PARTIA L-A DM ISSION IM PU LSE BLA D IN G FOR
H IG H-CAPACITY TO P T U R B IN ES
R e v i e w o f G e n e r a l P r o b l e m
Partial-admission impulse blading has been a standard element
of design over a long period. However, the industry has experienced
numerous troubles with such blading in the last few
years. It may well be asked, what has happened so quickly to
upset the design practice established by half a century of practical
turbine experience with partial-admission wheels The
answer is believed to be that the magnitudes of the suddenly applied
steam forces have been greatly increased beyond the
limits of the older practice, largely because of the introduction of
high-capacity top turbines at 3600 rpm and, at the same time, the
operating temperatures have increased to such an extent that
the physical properties of blading materials have been substantially
reduced.
The older forms of analysis employed in the design of partialadmission
impulse blading were based on static values of the
tangential steam driving forces, usually without allowances for
stress amplification due to the rapid rate of steam load application
and removal, or the effect of the application of the steam driving
load in successive revolutions under conditions of resonance.
In the earlier partial-admission wheels, the tangential steam
driving forces were of much lower magnitude, and the blade
heights were also less; hence, the effect of the repeated application
of the rapidly applied steam loads did not generally set
up destructive stresses. In such machines, the errors introduced
by neglect of the stress-amplification factor did not usually cause
trouble because of the small magnitudes of the bending stresses.
Where troubles developed in the older practice, due to exceeding
the safe stress limits accidentally, they were usually corrected
by the installation of stronger blades, sometimes without very
complete knowledge of the exact cause of failure.
Blade developments for high-capacity top turbines have led
to an extensive investigation of the performance of partial-admission
impulse blading, with the result that it is now possible
to predict with reasonable accuracy the maximum possible
stresses in such blading.
S t r e s s A m p l if i c a t i o n D u e t o E f f e c t o f S u c c e s s iv e I m p u l s e s
In the following discussion, the successive steps in the development
of the theory of dynamic loading will be reviewed. A
cursory review of the elementary principles involved because of
their importance w-ill also be included.
Fig. 1 illustrates the elementary principle of suddenly applied
loads. The spring in the diagram represents the elastic-spring
scale of a given blade. The mass of the platform can be taken
to represent the mass of the blade. The weight W represents the
tangential steam driving force.
At the instant the load W is applied, the full magnitude of the
force acts as an accelerating force. After an arbitrary deflection
I is reached, the force W is still acting at the constant value established
by its sudden application. There has been built up, however,
at deflection I, an opposing spring force which is equal to
the deflection I times the spring scale P. In other words, at the
arbitrary deflection I, the net accelerating force is the difference
between the steam force W and the opposing spring force PI.
The elementary theory shows that the triangular area having a
horizontal base W and an altitude Li represents, to some scale,
the energy that is imparted to the blade, which must appear in the
mid-position as kinetic energy. Note that at the static deflection
L\ the force W exactly equals the opposing spring force, so
that there is no acceleration at the mid-position.
If no damping is present and if the spring is assumed to have a
straight-line characteristic, it is evident that the weight will
overtravel a distance Li, which is equal to Li, in order that the
kinetic energy shall be converted into strain energy in the spring.
Therefore, the stress in the spring at the maximum extension is
twice the stress at the equilibrium position.

692 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
Fig. 2 shows a model which illustrates the same principle.
Without the small weight, the pivoted spring-loaded beam will
come to rest in the mid-position. This represents the position
of a blade under the action of centrifugal force only. A slow
application or removal of the steam driving force will cause the
blade to fluctuate between the position of rest and the normal
static deflection. This type of loading, the maximum value of which
is indicated by the first index mark below the center, represents
the operating conditions found in full-admission blades. If the
weight representing the steam force is engaged with the hook
blit just out of contact with it, and is suddenly released, the model
blade will deflect a greater amount than represented by the static
deflection. In the limit case, as shown by the elementary theory,
the maximum deflection will be twice the static value and will be
represented by the lower index mark on the model.
Let it be assumed that there is no internal friction in the
model and that the sudden application of the weight and the associated
vibratory motion continues with no reduction in the
amplitude. If the weight is suddenly removed while the blade
is at the lower extreme of its motion, the model blade will then
vibrate at double the former amplitude. The motion will continue
from a positive position corresponding to a tension in the
inlet edges of twice the static value to an equal stress in the
negative direction. In other words, with a sudden single application
of load under conditions of no damping, followed by a sudden
removal of load, timed to produce the maximum amplitude,
the stress range in the inlet edges will be 4 times that indicated by
the static value.
These statements indicate only the beginning of the process.
Let it be assumed, in the elementary case, that the load has been
suddenly applied and suddenly removed and that the blade is
now vibrating through the full amplitude indicated by theory.
If, now, the blade enters the admission arc a second time with
the motion in the right relation to the action of the steam force
imposed by the nozzle, it is obvious that the amplitude will be
further increased. The release of the steam force at the correct
time will yet further increase the amplitude of vibration. The
motion of the blade is thus increased in successive revolutions
until, if there is no internal damping of the blade material or
other friction effects, the blade would soon vibrate at a destructive
amplitude. This condition of timed application and release
of the steam forces is called primary resonance.
It is necessary to define primary and secondary resonance of
partial-admission blading. Primary resonance is considered to
be that state of motion in which the application and release of
the steam driving force occurs at the phase relations to the blade
motion which promote the maximum amplitude of vibration.
This type of resonance is probably unavoidable in steam-turbine
practice. It is, therefore, necessary to proportion the design of
partial-admission blading so that primary resonance can occur
and that the maximum amplitude set up by this condition will
not be such as to produce harmful stresses. If, under a given
operating condition, the application and release of the steam
driving forces are timed to cancel part of the energy input and
thus produce small amplitudes of vibration, which is quite possible,
this cannot be expected to be the case for all operating temperatures
or speeds. Hence, it is necessary to design the blading
to run under a condition of primary resonance.
Secondary resonance is the condition which exists when the
natural frequency of vibration of the blades is equal to the
frequency of the impressed force from the nozzles. This condition
is one which can be avoided by design, by proportioning
the blades to secure the required mass-stiffness ratio. The
periodic variation of the steam driving force, as the blades pass
through the admission arc, may set up excessive motion if secondary
resonance is not avoided.
In the practical case, in which internal damping is present, a
maximum amplitude of the vibratory motion is reached in which
the energy dissipated by damping equals the energy input from
the intermittent steam forces. This criterion of the conditions
regulating the stress-producing forces is capable of theoretical
analysis and permits the determination of blade proportions which
will survive such service.
It is of interest to note, in passing, that the kinetic energy imparted
to an initially nonvibrating blade at the end of the admission
arc is 4 times the amount imparted to the blade at the
inlet end of the arc. This statement is correct only for one application
of the load to a nonvibrating blade and one properly
timed release of the tangential driving force in which the application
and release each take place in zero time.
in this example is assumed to have no damping.
F i g . 3
The system
D ia g r a m S h o w in g E f f e c t o f A c t u a l T im e o f L o a d in g
o n B l a d e M o t io n
The foregoing discussion is based on suddenly applied loads.
In an impulse wheel for a 50,000-kw top turbine, a standard design
employs 66 blades. Each impulse blade will, therefore, receive
its loading in approximately Yea of Vso sec, or in approximately
V4000 sec. Fig. 3 shows the substantial modification made
in the elementary theory by the actual application of the steam
load in Vmoo sec instead of in zero time.
As the blade enters the admission arc, the steam driving load
starts at zero instead of at the maximum value which is established
in the case of the suddenly applied load. The tangential
driving force on the blade builds up as shown by the curve of
steam force on the chart, Fig. 3. At the left of the zero-force line,
the straight-line relation of spring force to deflection is shown.
The accelerating force at any deflection is the difference between
the force exerted by the steam and the opposing elastic force exerted
by the blade.
The actual kinetic energy imparted to the blade is proportional
to the shaded area at the right of the zero-force line. The
energy in this case is about */t of that imparted to the blade by a
suddenly applied load. On this basis, the stress-amplification
factor for a single application of the driving force at a uniform
rate in V«oo sec would be 1.35 instead of the factor 2, which is
found in the case of the suddenly applied load. This analysis

ALLEN—STEAM-TURBINE BLADING 693
shows the substantial influence on the energy input to a blade
under conditions of rapid loading in the short time interval of
Viooo sec, during which the full tangential driving load is built up.
As indicated, the loading of such a blade in V«»o sec results in 80
per cent less energy being imparted to the blade to start it vibrating
than in the case of a suddenly applied load.
In the consideration of partial-admission impulse-blade stresses,
the first step taken in the development of a rational design procedure
was the principle of suddenly applied loads described in
the preceding paragraphs. The second step was a process which
included consideration of the actual time in which the blade received
its load, the maximum amplitude of vibration which could
be built up by successive impulses received and removed under
conditions of primary resonance, and the principle that the energy
input to the blade from the nozzle must be absorbed by the
total damping of the system.
The third advance in the analytical reasoning on partial-admission
impulse-blade design is divided into three parts. The
theoretical work involved is the development of H. D. Emmert.3
The three divisions are as follows:
a Theoretical solution of the effect of suddenly applied and
suddenly removed steam forces on a vibrating blade, including
damping.
b Theoretical development which includes correct consideration
of the effect of the actual rate at which the steam load is applied
and removed on a vibrating blade.
c Development of the complete solution of the practical problem
in which the actual rate and time of steam loading of a
vibrating blade is taken into account; allowance is made for the
variation of steam driving force on the blade as it passes through
the arc of admission; and the correct provision is made for the
-ffect of the actual rate and time at which the steam load is removed.
The motion of elastically supported masses such as blades
under the action of intermittent forces is a type of problem which
F io . 4
F u n d a m e n t a l E q u a t io n o f M o t io n o f E l a s t ic B o d ie s
has been fundamentally developed on a theoretical basis. The
solution of the problem starts with the equation of motion, which,
in simple language, is an equation the terms of which represent
the various forces acting on the body.
The equation of motion as applied to this problem, referring
to Fig. 4, may be expressed as
t = time
01 — circular frequency of impressed force
k = spring constant
r = damping constant
a, 6, and c — steam-force constants
In the first term, m represents the equivalent mass of the vibrating
body. The remaining part of the term is the second derivative
of the displacement with respect to time. This quantity represents
the acceleration of the body at any point in its path. In
other words, the first term represents the inertia force, which is
equal to the mass times the acceleration at the point considered.
The second term represents the damping force. The damping
constant r is multiplied by the first derivative of the displacement
with respect to time, this quantity representing the velocity. In
other words, the damping force equals the product of the damping
constant and the velocity.
This assumption represents the conventional procedure and is
based upon the damping forces varying directly as the velocity
of displacement. The actual damping forces of materials of
construction may vary from this relationship. Such variation
may be taken into account by experimental determination of the
damping curve, and by using the value of the damping constant
which is applicable for the actual working condition.
The third term represents the spring force. The term k is the
spring scale and y the displacement.
The right-hand member of the equation represents the action
of the steam forces on the blade.
The motion of a blade in the inactive part of the circumference
is such that there are no steam forces acting, hence each of the
constants a, b, and c of the three terms is equal to zero during
this part of the motion.
If there is a condition of suddenly applied loading in which the
steam load is assumed to reach its full value in zero time, the
constant a will have a finite value, while the constants b and e
will each be equal to zero.
In the case where the steam load is assumed to be applied as
a linear function of time, the constants a and c become zero, and
the right-hand member of the equation becomes bt.
If the steam force in the admission arc is assumed to have a
harmonic variation which is found to be substantially correct
in practice, the constant b of the right-hand member becomes zero,
and the complete right-hand member for this case becomes a + c
sin wt.
The solutions of this type of differential equation of motion are
capable of being evaluated by established methods.
I m p o r t a n c e o f I n t e r n a l D a m p in g i n B l a d e M a t e r ia l
A factor of importance in the practical application of the equations
cited is the damping constant. A convenient assumption
in damping problems in hard metals is that the damping forces
vary as the first power of the velocity. This assumption is relatively
accurate over short ranges of operating conditions. On
this basis, it can be shown that the amplitudes of consecutive
vibrations have a constant ratio. If this law were strictly true,
the ratio of the amplitude of vibration at a predetermined time
to the next consecutive amplitude would be the same as the ratio
of any two consecutive amplitudes. The natural logarithm of
the ratio of any one amplitude to the corresponding amplitude
of the next cycle, taken as the vibration decays, is called the
logarithmic decrement.
The damping constant given in the equation of motion bears
the following relation to the logarithmic decrement

694 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
m = mass
/ = frequency of vibration
5 — logarithmic decrement
Fig. 5 shows a series of damping curves, made by Prof. John
T. Norton of The Massachusetts Institute of Technology, on
various materials and reported by him in a paper.4 These curves
are determined by means of a torsion-pendulum apparatus, in
which the specimen is subjected to alternating torsional stress.
The curves are exhibited to show the great differences in the
forms of the damping curves of the different materials. Considerable
research is now under way in connection with the damping
capacity of blading steels.
embedded a small weight. While the idea is ingenious, it does
not appear necessary to go to this refinement and, furthermore,
there are design and manufacturing difficulties associated with
this principle.
The first application of the third step in design reasoning is
shown in Fig. 6. The length of the abscissa in each case represents
the time required for the wheel to make 1 revolution. In
the upper curve, the loading is increased from zero to the driving
value in zero time and at the end of the arc again reduced to zero
in zero time. The lower curve shows the motion of a blade after
there has been a sufficient number of revolutions to set up a
condition of energy stability. The blade enters the admission arc
from the left with an amplitude of vibration which is reduced from
the value existing at the end of the previous passage through the
admission arc by the damping. The sudden loading increases
the amplitude and the corresponding stress as shown. In calculating
this case, it is important to note the assumption that the
load was applied when the blade under consideration was at the
extreme point of the backswing. This, on the chart, refers to the
lowest point of the harmonic curve. This is considered a condition
of primary resonance. In like manner, it is assumed that
the load is released at the time the blade is at the extreme point
of the forward swing, which corresponds to the upper point on
the harmonic curve at the end of the admission arc. The curve
shows the effect of damping in the admission arc and in the idle
arc. The reduction of stress amplitude due to damping has been
greatly exaggerated for illustrative purposes. By assuming acute
primary resonance, as described, the motion builds up in successive
revolutions, until, in the steady state of energy stability,
the energy imparted to the vibrating blade is equal to that absorbed
by damping.
In the case of the 50,000-kw top-turbine impulse blade, for the
Northwest Station, which will be described later, application of
At present for design purposes, with the Cyclops alloy, a
logarithmic decrement of about 0.002 is assumed for blading
operating with a maximum skin stress up to 20,000 psi and a
temperature of 900 F.
Foppl and others have shown that the damping constant
varies with stress. Other investigators have shown that the
damping varies with frequency and with temperature. The
research now under way will determine more completely the
damping capacity of the various blading materials under operating
conditions.
Proposals have been made and studied for the introduction of
damping devices in the blades themselves, such as a quantity of
liquid or powdered material in a cavity in the blade in which is
4 “ Torsion-Pendulum Instrum ent for Measuring Internal Friction,”
by J. T. Norton, Review of Scientific Instruments, vol. 10, March,
1939, pp. 77-81.

ALLEN—STEAM-TURBINE BLADING 695
the process just reviewed shows that the maximum bending
stress at the end of the first cycle in the admission arc will be
about 10 times the static value.
The second application of the equation of motion is shown in
Fig. 7. In this diagram, the tangential driving force is assumed
to be built up at a uniform rate, which is also assumed for a certain
condition of resonance, namely, the beginning of the application
of load is at the time when the blade is at the extreme
point in the backswing which on the curve is at the lowest
point in the motion. For illustrative purposes it is assumed that
the load application is completed in */* cycle of blade motion; in
other words, at the time the blade reaches the extreme forward
point in the swing. This case, which has been adopted for
simplicity of illustrative purposes, is not quite that corresponding
to exact resonance. In the latter case, the load is applied when
the blade is in the position of zero deflection and starting on the
backswing. This particular solution involves the assumption
that the steam load is constant throughout the admission arc.
The solution further assumes that the release of the steam
load starts at the time the blade is in the extreme forward point
in the swing, which is at the top of the harmonic curve in Fig. 7.
The condition of exact resonance at the release of the load would
be that corresponding to the similar state at the point of loading.
The third application of the analysis is illustrated by Fig. 8.
In this illustration, the motion of the blade in the idle portion of
the arc is treated as a damped free vibration as in the other cases.
The application and release of the steam driving load is assumed
to vary with time, this making necessary the treatment of
the admission and release phases of the motion as separate calculations.
As in the preceding case, the chart shows the load
application beginning at the time the blade is in the extreme position
of the backswing and ending at the extreme forward limit
of motion. The release is taken in the reverse manner.
The steam force imposed on the blade in the admission arc is
assumed to vary in the harmonic relation indicated on the chart.
When this process is applied to the 50,000-kw Northwest turbine
blade, the solutions of the equations of motion show that
the bending-stress-amplification factor, due to the considerations
mentioned, is approximately 3.
Impact-tube tests have been made with standard nozzles in
which compressed air is used instead of steam. Fig. 9 shows a
summary of the relative results of one set of such tests with the
impact-tube-pressure measurements taken at various circumferential
positions on the mean diameter of the nozzle.
It is interesting to note that the variation of the pressures per
unit area is much greater for the set of readings taken at an axial
D IAGRAM IL L U S T R A T IN G R ED U C TIO N IN FO R C E P U L S A T IO N
F i g . 10 D ia g r a m S h o w in g E f f e c t o f N o zzle P it c h on
V a r ia t io n o f S t e a m F o r c e s On B l a d e s a t D if f e r e n t P o s it io n s
A l o n g N o zzle F ace
clearance of Vi6 in. from the outlet of the nozzle than in the case
of the readings taken at a clearance of s/i6 in. This curve indicates
the substantial reduction in the variation of the forces per unit
area which occurs when a greater axial distance is allowed between
the outlet edges of the nozzle vanes and the inlet edges of
the blades.
An important factor in the design is the relation of the pitch
of the nozzles to the pitch of the moving impulse blades. The
upper dotted curve Fig. 10, shows typical steam forces per unit
area for various positions along the mean diameter. If the
integrated effect of this pressure can be assumed as transmitted
directly to the blades with equal efficiency for all positions of the
blades, then, theoretically, if the nozzles have the same circumferential
pitch as the blades, the tangential driving force imposed
on the impulse blades will be constant throughout the admission
arc. If, however, the transfer of steam driving load to the blades
does not occur with the same blade efficiency for all relative
positions of the blades and nozzles, then there will be some periodic
variation in driving force which is believed to be the case in
practice.
If the other extreme is considered, in which the blades are considered
as extremely narrow radial passages, then the driving
force on the blades will vary in substantially the same manner as
the steam force from the nozzle.

696 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
In order to design blades of relatively high frequency, a large
circumferential pitch must be used. In the case of the 50,000-kw
Northwest turbine blades, this pitch is 1.6 in. on the mean diameter.
The Northwest impulse blades are calculated to have a
running frequency of 7100 vibrations per sec. The nozzles have
been spaced on a pitch of 3.484 in., which gives a frequency of the
impressed force of 1837 per sec. The curve, Fig. 10, shows that
the theoretical variation of force on the blade is about 0.69 of the
variation of the equivalent nozzle steam force. There is, however,
a considerable margin between the 7100 per sec frequency of the
blade and the 1837 per sec frequency of the impressed force. The
magnification of motion from secondary resonance is relatively
small under these conditions.
In concluding this section on the stresses in partial-admission
blading, it can be stated that the investigations briefly reviewed
have shown that it is possible to calculate the effects of rapidly
applied loads on partial-admission blades. The only safe ground
for progress in matters of this kind is the establishment of a correct
background of fundamentals. It is believed that this has
been done and that the material constants are sufficiently well
known to permit the prediction of safe proportions for this type
of blading.
The importance of having the natural frequency of the blade
remote from the frequency of the variable impressed force from
the nozzle should be further stressed. Theoretically, if the action
of the steam on the moving blades occurs at the same efficiency
for all relative positions of blades and nozzles, then blades and
nozzles of the same circumferential pitch can be used to avoid
a periodic pulsation in the steam driving force. Actually, some
variation in the efficiency must be expected, as each blade moves
in and out of the flows of the various nozzles; hence, there will
be some periodic variation in the steam driving force on the blades.
If unsuitable design proportions are selected so that the frequency
of the impressed force on the blades is equal to their natural frequency,
then serious and destructive resonance may take place.
The best procedure is to design the blades so that the natural
frequency is substantially higher than the frequency of the impressed
force. The higher the frequency of the blades, the greater
the amount of energy absorbed by damping in each revolution.
The total damping will depend, to some extent, on the variation
of the damping constant with stress.
The general procedure for determining the stresses in partialadmission
blades can be outlined as follows:
a Determination of natural frequency of blade system including
correction for rotation and stiffening effect of shrouding.
b Determination of the frequency of the impressed force exerted
by the steam flow from the nozzle. An equivalent-steamforce
curve is assumed of simple harmonic form, the amplitude of
which must be based on research data.
c Design of blade to place natural frequency of assembled
structure substantially above the frequency of the impressed
force from the nozzles. Secondary resonance is thus avoided.
d The motion of the blade is calculated and plotted in curve
form for the condition of energy stability, in which the energy
imparted to a blade during one passage through the nozzle group
is dissipated by the total damping of the system in the course
of one revolution.
e Care must be taken to determine the most unfavorable phase
relations between the free motion of the blade and the time of
application and release of the impressed force. The phase relations
that will give the maximum amplitude of vibration must be
determined. In other words, the blades are designed to withstand
primary resonance.
f The correct procedure must be adopted properly to take
into account the actual rate of application of the steam load and
the corresponding rate of unloading.
g There must be reliable data available on the damping characteristics
of the blade material for the operating conditions under
consideration.
h The ratio of the maximum amplitude for the condition of
energy stability to the static deflection of the blade under the
action of the steam driving force taken as a steady value gives
the amplification factor to be applied to the bending stress calculated
on the static basis.
i The centrifugal stress must be added to the amplified bending
stress to determine the maximum resultant stress to which
the blade is subjected. All stress concentration factors must
be included.
C e n t r if u g a l S t a b il it y
There is good reason to believe that the initial tightness of
high-temperature blading in the spindle grooves is not maintained
for a long period of time at normal speed and full operating temperature.
This comes about partly due to initial creep in which
the minute irregularities of the machined surfaces are crushed
until a good bearing surface is secured and partly due to elastic
deformation and to the reduction in elastic modulus at high temperature.
It is believed that all manufacturers produce excellent
initial fits of the blade roots in the spindle grooves. At full speed
and temperature, however, the close fit is believed to be lost for
the reasons cited. If it be assumed that the blade root is only
slightly loose at full speed and full operating temperature, conditions
may arise, in which, in partial-admission wheels, the overturning
moment set up by the suddenly applied steam load may
exceed the stabilizing moment of the centrifugal force on the
blade. If this is true, a single blade, or even a group of two or
three blades which are joined together by a single shroud, may
jump in the groove every time the admission arc is passed. If
the blades jump or chatter in their grooves every time the steam
forces are imposed on them, local stresses of sufficient magnitude
to cause failure may be set up in the blade roots. It is, therefore,
necessary to establish a safe ratio between the stabilizing moment
created by the centrifugal force and the overturning moment set
up by the suddenly applied steam load. Experience indicates
that the ratio of the centrifugal moment to the steam driving
moment about the base of the blade should, in most cases, be
greater than 3 for a single blade, or greater than 6 for a pair of
blades tightly shrouded together. In general, it is believed necessary
that the stability factor shall be defined as the quotient of
the centrifugal stabilizing moment divided by the amplified steam
bending moment and that this ratio shall exceed 1.5.
Fig. 11 illustrates the principle of centrifugal stability. The
centrifugal moment is the product of the centrifugal force Fc and
the moment arm KL. The steam moment is the product of the

ALLEN—STEAM-TURBINE BLADING 697
modified by welding the shrouding in groups of two blades.
Fig. 13 shows a group of standard impulse blades for a tworow
wheel of the type used for the high-pressure element of 3600-
rpm condensing turbines up to about 75,000 kw or for 3600-rpm
top-turbine service up to about 25,000 kw. One of the important
features is the adoption of radial planes on the fronts and backs
of the root and shroud sections of the moving blades. In the
short first-row blades, the center portion of the root and shroud
on each side is in a plane passing through the axis of the spindle.
The overhanging edges have corresponding projections on the
root and shroud to reinforce them. The cooperating fit between
pairs of blades is on the central radial-plane surfaces. Note
that between the unwelded shroud sections, a clearance of approximately
0.005 in. is allowed during the cold assembly.
F i q . 14
A s s e m b l e d Sin g l e - R o w I m p u l s e W h e e l
F i q . 13
I n d iv id u a l I m p u l s e B l a d e s; D e s M o in e s
driving force F, and the moment arm L\. The stability factor
without amplification is, therefore, expressed as FeK L/F ,Li.
L im i t i n g C a p a c it y o f C i r c u m f e r e n t ia l - G r o o v e B l a d in g
Fig. 12 shows the high-pressure impulse wheel for the 35,000-
kw 3600-rpm condensing turbine at Des Moines, Iowa. The
first-row blades are IV 2 in- wide, and the steam-port height is
1 in. The steam load is 238 lb per blade at the '/a load point
on the turbine. This condition exists when the first two nozzle
blocks, to which steam is admitted simultaneously, are
operated at full inlet pressure. These blades are of Cyclops 17-A
alloy, with the integral shrouding welded together in groups of
two blades. The second-row impulse blades are shown welded
in groups of three. These second-row blades have since been
The limiting of the number of blades per group to two for hightemperature
service is believed important, as severe bending
stresses may be set up if a greater number of blades are joined
together. These stresses may be caused by the more rapid heating
and cooling of the blades with respect to the massive spindle.
Fig. 14 shows a single-row impulse wheel with the circumferential-groove-type
blading. This impulse wheel represents the
standard high-pressure element of large Allis-Chalmers highpressure
condensing turbines.
A x i a l -S l o t B l a d in g f o r H ig h e s t -C a p a c it y P a r t ia l -A d m is s io n
I m p u l s e W h e e l s
For maximum-capacity top-turbine service, the design analysis
briefly described calls for extremely wide blades. Such blades
have a large circumferential pitch, which makes it undesirable to
leave out a blade for assembly purposes. The centrifugal force
of one of these large blades is such that a practical locking device,
whereby the last blade can be secured in a circumferential groove
of the conventional design does not appear available. An extensive
survey has indicated the necessity for roots of the axialslot
type. Fig. 15 shows the impulse-wheel assembly for the
50,000-kw Northwest top turbine. The blades are inserted in
axial slots which are milled and broached in the spindle body. In
this way, the full complement of blades is inserted in the wheel
without the need of a special entry slot or locking blade.

698 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
F ig . 15
S in g l e - R o w I m p u l s e W h e e l 5 0 ,0 0 0 -K w T o p T u r b in e
In this scheme, the blades are arranged to be assembled in the
axial slots by entering the complete ring of 66 blades in the spindle
at the same time and pressing them home in a single operation.
This is necessary because the form of the blade section requires
that the edge portions overlap which, therefore, prevents individual
assembly. For this purpose, each blade is first tried in its
groove independently; each blade is then tried in its groove with
the adjacent blade on one side and a third time with the adjacent
blade on the other side, and so on around the circumference. By
this means, it can be determined that there is no mechanical interference
which will prevent complete assembly in one movement.
Figs. 16 and 17 show a model of the Northwest impulse blade.
These illustrations show the design of the serrated root and the
interlocking shrouding which requires the mass assembly of the
complete row of blades.
Note that the axial width of this blade is 3 in. through the port
section. The height of the steam port is l '/ s in. The mean
diameter is 34 in., as shown by Fig. 15. The shroud sections
are welded in pairs after assembly. The 4eep radial rib on the
F i g . 16 M o d e l o r I m p u l s e B l a d e f o r 5 0 ,0 0 0 -K w T o p T u r b in e F ig . 17 M o d e l o f I m p u l s e B l a d e f o r 5 0 ,0 0 0 -K w T o p T u r b in e

ALLEN—STEAM-TURBINE BLADING 699
integral shroud is for the purpose of increasing the strength of
the welded bonds between pairs of blades. The blades are
milled from Cyclops steel bars. The maximum steam load is
350 lb per blade.
Effect of By-Passing the Impulse Wheel on Blade
Loading
Some cases of trouble have been reported with top-turbine impulse
blades in machines in which all of the steam was supplied
through nozzles which acted on the first wheel. In such machines,
the first group of nozzles, with which the highest blade loading
occurs, occupied a relatively small arc of the circumference. This
in turn required relatively high blades with a correspondingly
greater length of moment arm on which the steam driving force
acted. The small admission arc for the first nozzle group resulted
in a correspondingly high driving force per blade. The combination
of these factors resulted in extremely high stresses in some
instances.
The latest top-turbine practice is to use the impulse wheel up
to about 3/t load, by-passing this element for full load, and further
by-passing some of the full-admission stages for maximum load.
By this means and by using only two nozzle groups for the impulse
wheel for the high-capacity machines, the first group occupying
approximately one half the circumference of the wheel, the
force per blade is substantially reduced, and, at the same time,
the moment arm on which the steam force acts is also reduced
greatly below the values found in former practice.
Effect of Avoiding Steam Undeeexpansion at the Outlet
of the F irst N ozzle Group
In maximum-capacity top-turbine design, it is an accepted
practice to design all of the first-stage nozzles of the nonexpanding
type,; in other words, the steam passages are designed to expand
the steam down to the sound velocity. In certain early machines
where all of the steam flow passed through the impulse-wheel
nozzles, the quarter-load-nozzle group was allowed to expand
the steam from 1200 psi down to something considerably less
than the critical pressure. This procedure resulted in a large
pressure drop through the blading and, in the case of some tworow
impulse wheels, a substantial upsetting of the load distribution
between the first and second moving rows of blades. With
maximum-capacity wheels, the present practice is to arrange the
first nozzle group to be fully open at the time that the pressure
in the first stage is reached which corresponds to the sound velocity
at the nozzle outlet. This practice aids in the reduction of
stresses.
Effect of Edge T hickness on F low Pattern Between
N ozzle and Impulse-Blade E dges
It is well known that the maximum efficiency in high-velocity
impulse blading requires thin blade and nozzle-vane edges. One
concomitant of 1200 psi initial pressure and 925 F initial temperature
is thick nozzle-vane-outlet edges and thick impulse-blade
edges; at least the edges are relatively thicker than found in
turbines of a few years ago. The increased edge thicknesses
necessary in high-pressure high-temperature machines have had
a small but detrimental effect on the efficiency. The steam-flow
pattern opposite each nozzle outlet is also adversely affected to the
extent that the alternating forces imposed on the blades as they
pass from the stream issuing from one nozzle outlet to the next is
believed to be the means whereby the amplitude of vibration of
the impulse blades is increased over what it would be with nozzles
and blades having thin edges.
Effect of Increasing Axial D istance Between Outlet
Edges of N ozzle Vanes and Inlet of Impulse Blades
In the interest of reducing as much as possible the variation in
loading on the impulse blades as they pass through the successive
flow patterns of individual nozzles, the axial distance between the
outlet edges of the nozzle vanes and the inlet edges of the impulse
blades is increased to about 9/ia in. in the case of the 50,000-
kw Northwest top turbine. Taken parallel to the absolute direction
of steam flow, the actual distance between the nozzle-vane
edges and the blade edges is about 2 in. This is expected to result
in some equalizing of the steam-flow pattern at the inlets to
the moving blades. It is believed that by this means less total
energy will be imparted to each blade in the form of vibratory
motion as it passes through the successive flow patterns of the
individual nozzles.
D E T E R M IN A T IO N OF STRESSES IN FULL-ADM ISSION
BLADING
Centrifugal Stresses
Centrifugal stresses in full-admission blading present no particular
problem as the determination of forces and areas may be
simply calculated from the drawing dimensions.
B ending Stresses
Bending stresses are calculated from the independent steam
forces exerted on each blade, and a knowledge of the section
properties of the blade section considered as a beam. The
tangential driving forces and the axial forces resulting from pressure
differences over the blade rows are independently considered.
R ipple Effect
Full-admission blades, especially in high-capacity high-temperature
machines, in which relatively thick blade edges must be
used, are expected to undergo some periodic variation in the driving
forces. This ripple effect probably does not exceed 10 or 15
per cent of the steady driving force. In high-temperature machines,
the ripple effect is taken as about 25 per cent of the steady
tangential driving force.
Shrouding Correction
The effect of shrouding varies considerably over the blade
path. In high-pressure blading, the effect of the shrouding on
stress is considerable, although the normal bending-stress reduction
may be between 25 and 60 per cent. The effect of the
shrouding on bending stress is complex, the calculations on stress
reduction being checked by experimental procedure.
Stress Amplification D ue to Vibration
The high-pressure and intermediate-pressure blading, in which
the natural frequencies of the blades are well over 5 times the
rotational frequency, are assumed to be sufficiently remote from
a resonant condition arising from periodic steam forces or inertia
forces caused by the rotation of the spindle that the stresses are
not intensified thereby. In low-pressure blading, in which the
frequency of the assembled blades at full speed and at the operating
temperature is less than 4 times the rotational frequency, tuning
is resorted to in order to avoid resonance under operating
conditions. This tuning involves a calculation of each specific
blade group, which is followed by a check of the standing frequency
when the row is assembled. Correction for temperature
and speed is made by calculation. Changes of blade sections, of
the rate of taper, or of lashing or shrouding members are sometimes
necessary in order to avoid resonance.
Stress Intensification D ue to F illet at Base of Port
Section
Milled blading with integral roots is designed with generous
radii at the junctions of the port sections and the roots, so that

700 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
with integral roots. Fig. 19 shows the large size of high-capacity
blades of this type. The nominal axial width of the last row
of blades, one of which is illustrated in the view of the model,
Fig. 19, is 21/, in. This blading is milled from bars of Cyclops
17-A alloy. Fig. 20 shows some reaction blades of the milled type
for top-turbine service.
D e s i g n o f S h r o u d in g
Referring again to Fig. 18, it will be seen that the shrouding on
the stationary blades is made up of two independent shroud sec-
F ig . 20
M il l e d R e a c t io n B l a d e s f o r T o p T u r b in e
F ig . 19
M o d e l B l a d e, L a st R o w 5 0 ,0 0 0 -K w T o p T u r b in e
the stress-intensification factor can usually be taken at 1.5.
For the intermediate blading with cast foundations, the bending-stress-intensification
factor, due to the form of the junction
between the rolled blade and the cast base, is taken at 2.
HIGH-PRESSURE REACTION BLADING
B l a d e s W i t h I n t e g r a l R o o t s M il l e d F r o m B a r S t o c k
Fig. 18 shows the blading layout for a 50,000-kw top turbine.
In this example, all of the blading is of the individual milled type
tions secured in place by welding. There are three radial-seal
strips for each blade row.
There is a groove in the spindle to contain the center-seal strip
and thereby provide interruption to the leakage flow. The inner
axial-seal strip formerly employed is absent in this construction
This is done to avoid the disturbance set up by the high-velocity
leakage flow cutting at right angles to the main flow through the
blade path.
The shrouding on the spindle blading includes four separate
seal elements; the first is of the normal end-tightened type; the
remaining three are of the radial type as shown. The shrouding
is of welded 19-9 chrome-nickel steel.
Note that the stationary blades of such high-capacity turbines
are also of the individual milled type.
W e l d i n g o f S h r o u d in g
The welding of shrouding has been adopted after considerable
research as being the best means for the attachment of such

ALLEN—STEAM-TURBINE BLADING 701
Fig. 22 shows the welded shrouding in the 30,000-kw top turbine
at Fisk Station in Chicago.
Blades Welded in Small Groups
It is important that the shrouding on high-pressure reaction
blading include a relatively small number of blades per segment.
The bending stresses introduced by rapid expansion and contraction
of the blading under conditions of varying temperature may
result in excessive stresses and lead to ultimate blade failures unless
short segments are used. Fig. 22 shows that the reaction
F ig . 23
A s s e m b l y D r a w in g o f S e g m e n t a l I n t e r m e d ia t e
B l a d in g W it h W e l d e d S h r o u d in g
F io . 21
W e l d in g o f S h r o u d in g on S e g m e n t a l B l a d in g
F ig . 24
I n t e r m e d ia t e -B l a d in g S e g m e n t s W it h W e l d e d
S h r o u d in g
blades in the Fisk turbine have shrouding which contains from
four to six blades per segment. It has already been pointed out
that the impulse-blade practice is to secure two blades together
for the highest-temperature service.
IN T E R M E D IA T E BLADING
F ig . 22
W e l d e d R e a c t io n-B lade S h r o u d in g in F is k T u r b in e
members. The stresses set up by riveting are avoided. By the
use of austenitie materials, air-hardening does not occur. Fig. 21
shows a group of blades with a section of shrouding ready for
welding. A welded shroud section is also shown.
B lades Rolled to F inal Sections
For the intermediate-pressure blading, sections are used which
are finished to the final shapes by rolling. These blade sections
are of Cyclops 17-A alloy. The material is given a final stressrelieving
heat at 1425 F after the rolling operation. Before using
the sections, the inlet and outlet edges are carefully polished to
remove minute cracks which might have been introduced during
the rolling. Fig. 23 shows in section the standard arrangement of
intermediate blading. The welded shrouding and the design of
the seal strips are shown. Fig. 24 shows a group of intermediate
blades of the cast-in type with welded shrouding.

702 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
The study of the rolling of the blade sections and the improvements
made at the steel mill have been previously described.
Shrouding D esign
Blade sections, after being cut to length and after the inlet and
outlet edges are polished, are assembled in a special fixture, after
which the 19-9 shrouding is welded in place.
Cast F oundation of Everbrite or Cast Iron
After the shrouding is welded in place, the segmental-blade
assembly is then placed in a core box, after which the blades and
shrouding are encased in core sand, with the exception of the
ends of the blades that are to be cast in the foundation. The
blades in their assembled core are placed in a suitable mold and
the foundation of Everbrite or cast iron poured. Great care is
necessary to maintain clean blade roots. It is also necessary to
keep the pouring temperature of the base material within close
limits in order to secure a good bond with the blade bases without
washing the blade-section material away. Optical-pyrometer
observations are made on the metal poured in each mold.
Stress-Relieving of Assembled
Sections
Cast F oundation-R ing
After the foundations are poured, they are then finished on all
surfaces, except at the bottom of the steam port, by machining.
F i g . 25
M il l e d L o w - P r e s s u r e B l a d in g
After machining the foundations, the segments are placed in
a gas-charged muffle, where they are annealed at 1425 F to remove
stresses introduced by the cooling of the foundation metal.
By this process, completely finished segmental blading is manufactured
for the intermediate portion of the blade path. The blading
is practically free from fabrication stresses and in an excellent
condition to resist corrosion.
M illed Low-P ressure R eaction B lading
When the blade annulus referred to 3600 rpm exceeds approximately
9.2 sq ft, the blades are milled from bar stock of Cyclops
alloy with integral roots, as shown in Fig. 25. Care is exercised
in the mill operations to secure a grain size not exceeding 20 per
sq in. at 100 diam magnification. The bar stock is stress-relieved
at 1425 F after the final rolling.
The milled Cyclops low-pressure blading is reinforced by
shrouding of 19-9 chrome-nickel steel which is welded in place
as described for the high-pressure blading.
The use of lashing wires is avoided except in the case of lowpressure
blades where they are useful for tuning purposes. Lashing
wires are not used in blades where the natural frequency of
the assembled blade groups under operating conditions is more
than 4 times the rotational frequency.
When lashing wires are used in Cyclops blades, the blade is
drilled and the edges of the holes suitably rounded. A short
section of 19-9 chrome-nickel wire is welded in place in each
blade. After welding, the blade is stress-relieved at 1125 F.
When the blades are in place in the spindle, the ends of the individual
lashing wires are joined in groups by welding. In this
way the heating effect of the final-assembly welding does not
reach the body of the blade.
LOW -PRESSURE BLADIN G FOR M AXIM U M T IP SPEEDS
Limiting Capacity of Condensing M achines
At the present time, the limiting exhaust-blade practice is a
20-in. blade with a tip speed somewhat over the sound velocity
in steam at 29 in. vacuum. In Allis-Chalmers 3600-rpm practice,
20-in. blades are mounted with a tip diameter of 80.5 in. with a
corresponding tip speed of 1264 fps. Regarding the capacities
for which these blades can be used, this selection involves a commercial
consideration of the actual load requirements and operating
charges for a given system, which will in turn indicate the
amount of leaving loss which can be allowed. If operating conditions
of 1200 psi gage 925 F and 29 in. vacuum be assumed with
four stages of feed heating,
the leaving loss of a
double-flow low-pressure
end with 20-in. blades at
3600 rpm will be as given
in Table 1 (on the following
page) at the respective
capacities. These blades
are assumed to be mounted
with a tip diameter of 80.5
in. and with an opening coefficient
of 0.6.
The present outlook
does not indicate a radical
tendency toward an increase
of tip speeds. The
length of the Allis-Chalmers
blades described is
just under Vs of the mean
diameter. The form of the
passage at the tip section
must be a compromise
with some sacrifice in efficiency
in the interest of
practical construction. An
appreciable increase in
blade length with respect
to the mean diameter does
not seem probable at this
time.
The effect of water erosion
is sufficiently serious
F i g . 26 A x ia l -S l o t T y p e 1 8 -In .
3 6 0 0 -R pm L o w - P r e s s u r e B l a d e

ALLEN—STEAM-TURBINE BLADING 703
T A B L E 1
Leaving loss expressed as a per cent of total
adiabatic drop including corrections for
Rated load feed heating and underexpansion
40000 kw 1.9
50000 kw 2.9
60000 kw 4.0
at the present tip speeds to preclude any strong tendency for
increased tip speeds. The present trend of 3600-rpm construction
is, therefore, in the direction of multiple-cylinder low-pressure
elements where maximum-capacity units are considered.
The present limiting annular area for 3600-rpm low-pressure
blades is about 26.4 sq ft. This is the area corresponding to the
20-in. blades mounted with a tip diameter of 80.5 in. previously
referred to.
Fig. 26 shows an 18-in. blade of this type with a tip diameter of
76.5 in., with a corresponding annular area of 23 sq ft. Fig. 27
shows an assembled row of these blades in place in the spindle,
and Fig. 28 shows the axial-entry spindle slots in which the
blades are installed. The design of these blades is based on a
cross-sectional-area ratio of 6, i.e., the sectional area of the base
of the blade is 6 times the area at the tip. The ratio of mean
diameter to length is approximately 3, which means that there
is a substantial change in the contour of the steam passage from
the base to the tip.
For annular blade-outlet areas between 21 and 26.4 sq ft at
3600 rpm, the 13-chrome stainless iron is used because of its
greater strength at low temperature. This material is forged to
within about Vi« in. of the finished surfaces and then completely
machined to the final sizes.
Shrouding and lashing members are necessary in the last row
of blades to reduce bending stresses and to bring up the natural
frequency of vibration to the required value. A section of
shrouding, one blade pitch in length, is welded to the tip of each
low-pressure blade. The shrouding is of the 19-9 chrome-nickel
steel. A short stub end of 19-9 chrome-nickel lashing wire is
welded to each side of each blade, thereby eliminating the lash­
F io . 27
A s s e m b l y 1 8 -In . 3 6 0 0 -R pm L o w - P r e s s u r e B l a d e s
F i g . 2 8
S p i n d l e f o r A x i a l - S l o t B l a d e s

704 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
F iq . 29
F i g . 30
N o zzle E l e m e n t s
N o zzle E l e m e n t s
ing-wire hole. A generous coating of stellite is welded on the inlet
edge of each blade for about 2/s of the length from the tip to resist
water erosion. The stellite is fused to the blade with an acetylene
torch. The blade is preheated to about 900 F for these welding
operations.
After the welding is completed, the blade is heat-treated by
quenching in oil from 1750 F and drawing at 1125 F. In this way
the harmful effect of the air-hardening is greatly reduced.
After the blades are assembled in the spindle, the corresponding
shroud and lashing members are joined by welding. The
welding after assembly is entirely between non-air-hardening
materials. The local heating of the final welding operations does
not extend to the body of the blade.
The sectional shrouding is of cantilever form, each section being
secured to the blade by a generous welding bead. The extreme
overhanging ends of each cantilever are joined by light
welds after assembly. The overhanging cantilevers are designed
to be stable in the event of accidental breakage of the assembly
welds so that failure from centrifugal stress will not occur.
The two machines, equipped with 18-in. low-pressure blades
which operate at a tip speed of 1200 fps, are in their second year
of service and have thus far disclosed no troubles of any kind in
the low-pressure blading. Two rows of these blades are installed
in the low-pressure turbine of the 35,000-kw 3600-rpm turbine
at Des Moines and one row in the single-flow machine at Springfield,
111. From this experience, there is good reason to expect
that the 20-in. blades, which will be designed on the same engineering
principles and stress limits, will also operate satisfactorily.
N o z z l e D e s ig n
Figs. 29 and 30 show the type of nozzle elements used in highpressure
high-temperature machines. The nozzle ports are
machined in individual stainless-steel elements which, when
assembled, form the accurately machined steam ports. The
assembly drawing of a similar nozzle is shown in Fig. 31. Each
individual steel block in the group of central elements is defined
by radial end planes which are carefully fitted together. After

the central group of elements is assembled, the inner and outer
ring segments are fitted over tenons and welded in place.
A c k n o w l e d g m e n t s
The author wishes to thank the Allis-Chalmers Manufacturing
Company for permission to publish the blading-design data included
in this paper; Professor J. J. Ryan, of the University of
Minnesota, and Dr. J. T. Rettaliata, of the Allis-Chalmers steamturbine
department, for their suggestions during the early development
of the partial-admission impulse-blade reasoning;
Professor John T. Norton, of the Massachusetts Institute of
Technology, for suggestions in connection with damping phenomena
and for permission to use Fig. 5; Mr. H. D. Emmert,*
of the steam-turbine department of Allis-Chalmers, for his
assistance in preparing the mathematical solutions of the various
phases of the motion of vibrating blades; and to numerous other
members of the Allis-Chalmers organization for their help in
the preparation of this paper.
Discussion
R . P. K roon.* Air-hardening 13 per cent Cr steel and nonair-hardening
nickel-chrome austenitic steels have been described
as applied to turbine-blade service. Many years ago the
Westinghouse Company adopted 13 per cent Cr blade steel because
it was one of the few corrosion-resistant alloys adapted to
forging. More recently, the forging process has been abandoned
for manufacturing reasons on all except low-pressure blades.
We are, therefore, in a position to use any of several high-temperature
alloy steels suitable for machining.
A large number of such materials were made into blades for
experimental purposes and some have been used in turbines in the
field. Among these materials, the 13 per cent Cr steel has distinguished
itself because of its superior yield strength and its
metallurgical simplicity and uniformity.
We recognize the problem of air-hardening in 13 per cent Cr
steel as a result of shroud welding, but have used this process
for a number of years under proper control and with but slight
difficulty. I t has been found recently that the 13 per cent Cr is
unique in that it possesses many times the damping capacity of
any other high-temperature material known to the art. We continue
to regard the 13 per cent Cr stock as that having the most
desirable combination of qualities for turbine blading.
We cannot very well agree with the statement that blades
having a natural frequency well over 5 times the running frequency
are sufficiently remote from resonance so that the stresses
are not intensified thereby. I t is our experience that, even with
full admission, the resonance effect is important at frequencies
up to say 12 times the running speed. We are using statistical
data to insure a safe design for these medium-size blades.
By this time it should be well recognized that the impulse-blade
problem is industry-wide. It behooves us that this fact should
be admitted; the problem is there whether we do so or not.
The writer has been particularly interested to find that the
assumptions, on which the author has based his calculations, are
almost identical with our own original suppositions. However,
we have felt very strongly that quite a few of these assumptions
are based on such loose grounds as to make necessary the experimental
determination of the exact nature of the forces and
motions involved in actual operation.
1 These solutions in detailed form have been presented by H . D.
Em m ert in a “ M athem atical Section” of this paper, as originally
published by the Allis-Chalmers M anufacturing Company, prior to
the 1940 Semi-Annual Meeting of the Society. Copies are available
on request to th at Company.
8 Manager, Experim ental Engineering, Westinghouse Electric &
M anufacturing Company, South Philadelphia, Pa. Jun. A.S.M.E.
ALLEN—STEAM-TURBINE BLADING 705
We started a research program using optical equipment to
measure actually the blade deflections, first on a small pilot
turbine which was placed in operation 9 months ago, then on a
full-size Curtis turbine, operating a t 1250 lb 900 F. This work
was recently described.7 We are more and more convinced that
this investigation is of vital necessity to give those data which
cannot be derived theoretically.
There is, for example, the assumption that a blade receives its
load gradually as it travels one blade pitch. In the absence of
experimental data this assumption is probably as good as any
other. But one must not forget that actually the forces on the
blade are caused by the reactions of a very complex nonsteady
steam flow on either side of the blade. Actually, we find that it
may take a distance of several blade pitches before the blade
force steadies out.
Instead of a gradually increasing steam load which the author
has assumed, we often find that the first shock on the blade entering
the steam jet is in a direction opposite to the torque force.
This can be explained physically by the fact that, as the steam
first enters the blade passage, it tends to push the blades apart.
This pressure wave causes a negative rather than a positive torque
force on the trailing blade.
Our findings have also indicated that, in general, it is not
sufficient to limit consideration to vibration in the circumferential
direction only. We formerly believed that the lateral
vibrations of the blade were not important, but we now know
that, on an impulse-blade group, there are at least three natural
modes of vibration which should be considered. The first is a
vibration in the plane of rotation, the second is one in which the
blades vibrate perpendicularly to the plane of rotation, and the
third is a torsional motion of the group as a whole. Under certain
conditions, it is necessary to take into account all three
modes.
In our opinion, expanding the steam down to more than sound
velocity in nozzles of the nonexpanding type is quite feasible,
provided that this superacoustic range is explored experimentally
to establish safe limits. This we are doing in our developmental
turbine. We believe that, with this background, the additional
efficiency at low turbine loads can be made available without a
risk.
The statement is made in the paper that, with a gradually increasing
load, the maximum displacement of the blade will be
produced when the initial displacement is zero, as the load is applied.
T hat this is not so can be appreciated when one considers
the limiting case of suddenly applied load, for which we
know that maximum displacement is produced when the initial
velocity, and not the displacement, is zero at the instant the load
is applied.
The writer believes the author will be able to check his conclusions
that the phase angle for optimum displacement depends
upon a relation between loading time and natural period and
trusts he will see fit to revise his calculations to take this into
account.
G. B. W a r r e n .* The author has presented an interesting
analysis of the forces on the first-stage partial-arc-admission
impulse-blade element, which should have common application
in varying degrees to all types of turbines incorporating this
construction. Such analyses should be of great value.
The group of engineers with which the writer is associated
has followed a somewhat different method in its endeavors to
attain the same result, namely, a stable and reliable construction
7 “ Superposed Turbine Blade Research,” by F. T. Hague, Mechanical
Engineering, vol. 62, April, 1940, pp. 275-277.
8 Designing Engineer, Turbine Engineering D epartm ent, General
Electric Company, Schenectady, N. Y. Mem. A.S.M.E.

706 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
for the heavily loaded buckets of first-stage impulse elements.
Fortunately, we had as background nearly 35 years of experience
in building first-stage partial-admission impulse buckets,
which included several hundred rows, a few of which had given
trouble in service. We became convinced during the course of
this experience that the difficulties were due to vibrations produced
by the operating conditions. However, it has always been
felt that these vibrations and their consequent stresses were so
complex and of such varied form as to make any simple rational
analysis inexpedient. A statistical study of the conditions under
which such troubles occurred, and those under which they were
not present, indicated a rather definite function of the ratio between
the severity of the conditions and the strength of the bucket
and its attachment. Formulating this knowledge into design
practice was undertaken about 12 years ago with the result that,
except for a few isolated and special cases, difficulties with
these elements have practically disappeared. Recent tests appear
to indicate that no sacrifice in efficiency seems to result from the
use of the wider and, hence, stronger buckets.
The indications from the paper are that the author’s analysis
is leading him to conclusions of about the same type and magnitude
as our own.
As to the details shown, we differ in materials used, in attachments
of buckets to rotors, in method or shrouding the buckets,
and innumerable details of design and construction. The fact
that two or three turbine manufacturers differ in such things is
probably the greatest safeguard to the power industry. It insures
a minimum average of trouble, and the maximum possible
progress.
The design and construction technique shown for the last rows
of blading on limiting-capacity turbines is most interesting.
From the illustrations, it seems evident that the shroud band
and lashing members finally are reduced to one continuous
piece. Some additional discussion of this construction would be
of interest.
A. L. K i m b a l l .9 This paper brings to focus two things, i.e.,
the importance of avoiding a condition of resonance, wherein a
periodic exciting force is in step with one of the natural frequencies
of the elastic bucket system; it brings out the importance of
safeguarding possible dangerous resonant-vibration amplitudes
by having as much internal damping in the buckets as possible.
The dangerous effect of resonance is well shown from the authors’
Equation [19] ,10 by setting the exciting frequency equal to
the natural frequency so that a = oin. In this case U is seen
to be equal to zero and R reduces to the logarithmic decrement
$ divided by v or R = 8/ir. The last two terms of Equation [19]
then become
from the case of a simple linear oscillator,11 which alone gives us
the important part of the picture.
Note that, since the deflection given by the authors equals
0.000035 in. for a steady force of 350 lb at the tip of the bucket,
an equivalent force of amplitude 55,000 would be necessary to
produce the deflection at resonance given by Equation [1] of this
discussion, where the term amplitude in this case, equals the +
or — amplitude of the periodic force, the total range of which
from + to — is therefore 110,000 lb. This can be seen to produae
a maximum stress far beyond the fatigue limit when it is noted
that the total volume of this bucket is about 0.6 cu in. of material.
Thus, if it is assumed to be of rectangular shape, 1 in.
high, 1 in. wide, and 0.6 in. thick, the maximum vibration stress
at its base is over 500,000 psi.
The analysis in this paper gives an excellent idea of the importance
of impulse excitation of bucket vibration, and shows
that its intensity is always far below that arising from resonance
produced by even a moderate periodic force, which was taken
as 0.1 the average bucket load in the author’s case. Furthermore,
the analysis also shows that the damping in the buckets
is not an important factor in controlling impulse excitation.
The real reason for damping is to control the amplitudes which
might arise on resonance as given by Equation [2] of this discussion.
The writer personally dislikes the use of the terms “primary
resonance” and “secondary resonance.” The former represents
a condition of impulse excitation and is not a resonance at all
in the sense that the second type is, which is by itself the true
resonance.
A complicated elastic system like that of the bucket zone of a
turbine wheel may have several modes of high-frequency vibration
to each of which a resonant vibration will build up if in step
with some periodic exciting force, and some of which may be
dangerous. Furthermore, such modes of vibration are not readily
predicted by available methods of analysis, so the only sure way
is to find what they are from actual test.
Prevention of such failures therefore requires:
1 The definite location of the type and frequency of possible
modes of vibration from actual tests, which may be in the range
of known high-frequency exciting forces which arise during
turbine operation.
2 A thorough study of the damping of bucket materials
under operating conditions, the latter being an insurance against
danger from the former.
3 Adequate area of cross section for safe performance.
These methods of attack are now being applied by turbine
manufacturers who appreciate the danger only too well, and
steady progress is being made.
This is seen to be over 50 times as great as for the case chosen
by the authors which gave ym = 0.000106 in. Furthermore,
the first term of Equation [19] was omitted, which makes the
value of Equation [2] of this discussion conservative.
Note also that Equation [1] herewith may be directly derived
* Engineering General D epartm ent, General Electric Company,
Schenectady, N. Y. Fellow A.S.M.E.
10 This discussion refers not only to the paper proper by R . C.
Allen but also to a m athem atical section by H. D. Em m ert, which
was presented at the meeting. Space did not perm it publication of
this section in the Transactions, but copies of the m athem atical treatm
ent may be obtained upon request to the authors.
T. C. R a t h b o n e . 12 The authors10 are to be commended for the
presentation of this valuable and instructive paper and the research
behind it on blade-failure problems which have caused
some anxiety.
The writer is particularly interested in the reasoning leading to
the conclusion that, as some resonance in impulse blading is unavoidable,
the blades must be designed to withstand the most
unfavorable combination of conditions.
The frequency calculations are based on the assumption that
the blades remain tight in the root, that is, the node point is
fixed. I t was pointed out that looseness in the root with partial
admission permits a rocking impact or chatter which may well
11 “Vibration Dam ping Including Case of Solid Friction,” by A. L.
Kimball, Trans. A.S.M .E., vol. 51, 1929, paper APM-51-21, p. 230.
12 Chief Engineer, Turbine and M achinery Division, Fidelity and
Casualty Company of New York, New York, N. Y. Mem. A.S.M .E.

ALLEN—STEAM-TURBINE BLADING 707
F ig . 33
S c h e m a t ic R e p r e s e n t a t io n o f B l a d e-R o o t S y s t e m
have set up excessive repetitive stresses, explaining past failures.
By designing for a centrifugal-stability factor over the overturning
steam-force moment, the authors insure the nodal fixity at
the root necessary for the assumptions in the frequency calculations
which follow.
It will be interesting to point out here that, with any appreciable
looseness in the root, the natural frequency of the blade
becomes a function of the amplitude of vibration as well as of
its mass and elastic properties, and thus may undergo wide
variations. Fig. 32 (a) represents a blade with loose root fastening
in repose, and Fig. 32 (6) the position at one extremity of movement,
having rocked on one bearing shoulder until stopped by
the limiting clearance in the groove.
The system can be illustrated schematically by Fig. 33 (o),
where the blade is represented by a pendulum A rigidly attached
to a bracket B, serving as the root, and supported at the
two points C and D (root shoulders). Gravity (centrifugal
force) is pulling downward, attempting to hold the system in repose
while the pulsating steam force F is tending to rock the system
about the bearing point D.
If the pendulum is displaced slightly to one side and released,
it will oscillate or rock back and forth under the action of gravity
alone, at a frequency which increases rapidly as amplitude decreases,
precisely like a rubber ball dropped on the floor, until the
system comes to rest. So far, neither mass nor elasticity enters
the picture; only centrifugal force (gravity).
Assume now that the displacement is great enough for the
corner of the bracket or root to strike at E. First there will occur
an elastic deformation at E due to the impact, and also a flexural
bending of the pendulum in attempting to carry on in the leftward
movement. In addition to the steam and centrifugal
forces there, we have both bending and impactive forces at the
fillet P, with the stress-magnification factor due to impact
depending upon the “softness” at the contact points.
The restoring forces to complete the other half of the cycle now
depend on:
F i g . 35 I n s t a n t a n e o u s P o s it io n o f B l a d e M a k in g (a) 1 a n d (6) 2
C o m p l e t e V ib r a t io n s D u r in g T im e R e q u ir e d f o r 1 R e v o l u t io n
1 The elastic “rebound” at the impact areas.
2 The elastic and inertia properties of the pendulum system.
3 The gravity force acting through the clearance space E to
E '.
An intermediate impact also occurs in this simple system as
the rocking points C and D strike their seats.
The frequency of such a system depends not only upon the
mass and elastic characteristics of the blade itself, but also on the
impactive resilience, and finally on the extent of free movement
or amplitude limited by the clearances. The frequency is practically
indeterminate. The vibration of such a system can no
longer be expressed by linear differential equations, and hence is
called “nonlinear” vibration. Professor Timoshenko1* points
out that such systems may often nullify the effect of resonance
because, as resonant amplitudes build up, the change in amplitude
itself becomes a sort of inhibitor by altering the natural
frequency away from resonance.
There have been examples, however, where impactive interference
with free vibration caused a great increase in amplitude
and hence stresses. These cases seem to involve a critical condition
in which the elastic rebounds from impact somehow augment
13 “Vibration Problems in Engineering,” by S. Timoshenko, D.
Van N ostrand Company, New York, N. Y., 1937, pp. 117 and 145.

708 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
or aggravate critical vibration. Increased turbine vibration due
to impactive movement from excess clearance in turbine bearing
keys is a familiar example. This is further illustrated by the
model Fig. 34. In Fig. 34 (o) the cantilever or blade is shown
vibrating steadily in response to magnetic impulses; at (6), with
no change in the magnetic impulses, the vibration has been greatly
increased or “drawn out” by impactive interference near the
root.
I t is pointed out in the paper that, when the natural frequency
of the last-row low-pressure blades is less than 4 times the running
speed, timing is resorted to, to avoid resonance. Presumably,
the periodic forces from which this blading is to be protected
arise from torsional oscillations of the rotor or linear
vibration of the rotor due to unbalance, at the frequency of rotation,
rather than periodic steam forces.
Primary resonance occurs when the natural frequency of the
blade coincides with a torsional oscillation once per revolution,
but resonance with linear vibration from unbalance does not occur
until the natural frequency of the blade is twice the rotational
frequency. Therefore, a blade frequency of 4 times the speed of
rotation actually represents a multiple of only 2, as regards
fundamental resonance. Likewise, no resonant amplitude should
be built up when the blade frequency is just equal to linear vibration
at the frequency of rotation.
In Fig. 35 (a) are represented the instantaneous positions of a
blade making one complete vibration during the time required
for one revolution. I t is assumed that the blade vibration is
caused by the linear vibratory movement of the rotor due to unbalance,
acting on the blade through the root fastening. The
blade is shown in the neutral position at 1, and at the extreme
limit of its vibration in the leading direction at 2, after 90 deg rotation.
The blade passes through the neutral position again at
3, and reaches the opposite extreme at 4, in the trailing direction.
But the inertia forces on the blade from the rotor vibration,
which caused it to be displaced forward at 1 as shown in the first
part of the cycle, would again cause the blade to travel forward
at 3 during the second half of the rotor vibration cycle, instead
of the backward motion attempted by the free vibration of the
blade. In other words, the “return-stroke” exciting force opposes
the completion of the blade-vibration cycle, and should
discourage building up resonant amplitudes.
In Fig. 35 (6) is shown the situation where the natural frequency
of the blades is twice the running frequency. Here the
natural blade movements are in proper phase with the rotor
vibration to produce fundamental primary resonance.
Fortunately, as the author points out, there has been little
difficulty with fatigue failures of the last-row blades due to resonance.
Curiously, we had all been far more concerned over
the possibility of fatigue failures in these large blades than
in the rugged shorter blades which actually have given more
trouble.
A u t h o r ’s C l o s u r e
' With reference to Mr. Kroon’s discussion concerning the 13
per cent chrome blading steel, the author believes that all turbine
builders are in the same position as that assumed by Westinghouse,
namely, each is ready to use the material which he considers
the best for steam-turbine blading. That the 13 per cent
chrome steel has justified its adoption for this service is well
demonstrated by the number of turbine builders who make
successful use of this material. The author states in his description
that this alloy is the best material for low-pressure blading.
Many years of development and metallurgical research were
required before the 13 per cent chrome steel was produced as a
useful blading alloy. It was the author’s privilege to be associated
with Francis Hodgkinson and Norman Mochel during
the early development of this material by the Westinghouse
Company.
Paralleling the work of the Westinghouse and General Electric
Companies on the 13 per cent chrome steel, the Allis-Chalmers
Company has developed the austenitic Cyclops alloy described in
the paper. They, too, are in the position of being able to use the
blading steel that they consider the best for blading. The conclusion
that the Cyclops alloy offers advantages for all but the
longest rows of low-pressure blades is again a matter of many
years’ development and research which was only briefly indicated
in the paper. The author recognizes that each material has advantages
and disadvantages, so that a careful survey of the merits
of the two materials is necessary. The excellent service record of
the Cyclops alloy speaks well for its reliable qualities for blading
service.
With reference to the difference in damping capacity, this
property is still under investigation. It is necessary to point out
that, as stated by the author during the discussion of the paper
when presented, partial-admission impulse blading, such as
illustrated for the Northwest top turbine, is designed so that the
structure is safe with a logarithmic damping decrement as low
as 0.0002, including a stress-concentration factor of 2.0 due to the
fillet at the base of the blade. It is believed that the construction
described, in which due allowance is made for amplification of
the bending stress from primary and secondary resonance, is
based on logical reasoning and that safety is insured if the proportions
permit the operation with safe stresses with the low
damping decrement mentioned in this paragraph.
Mr. Kroon indicates that he is not in agreement with the
statement that full-admission blades having a natural frequency
over five times the running frequency are sufficiently remote
from resonance so that the stresses are not intensified thereby.
He also indicates that he has found the resonance effect to be important
at blade frequencies up to about 12 times the running
speed. The author’s experience is not in agreement with this,
although quite a number of cases have arisen in which this was
first thought to be the cause. Careful analysis of other effects,
such as pressure differences in the blade path or defects in the
mechanical construction, were in each instance sufficient to cause
the high resonance ratio to be abandoned as a reason for trouble.
Mr. Kroon points out that his company is using statistical data
to insure a safe design for blade structures. This practice is today
and, so far as the author is aware, always has been the u n d e rly in g
principle employed by all blading designers. Blading stresses
and calculation processes have always been checked and compared
and the constants evaluated by comparison with practical results
taken from field experience or laboratory research.
With respect to Mr. Kroon’s comments on the general character
of the impulse-blade development, the author agrees that
the assumption that a blade receives its load gradually as it
travels one blade pitch is much on the safe side as very elementary
considerations will serve to show. The author agrees that it may
take a distance of several blade pitches before the normal driving
force is established on the blade. The statement that, in the experimental
turbine built by Westinghouse, the first movement of
the blade entering the steam jet is opposite to the torque force is
very interesting. All of these effects appear to the author of a
nature that does not modify the general validity and safety of
the primary assumption of the paper with respect to the time of
loading.
The author is in agreement with Mr. Kroon, in that it is necessary
to consider the vibration of blade structures axially and
torsionally, as well as circumferentially. The study of impulseblade
structures in the paper is limited to the circumferential
motion, because, in wide-root blades of the type described, this
mode is the only one that need be considered. In long blades

ALLEN—STEAM-TURBINE BLADING 709
that may be shrouded and lashed together in groups, consideration
must be given to the other modes of vibration.
With respect to expansion ratios that correspond to more than
sound velocity, the author’s statement in this connection applied
specifically to high-capacity top-turbine design. In such machines,
it is the present practice to arrange the first group of
nozzles to be fully opened at the time the pressure in the first
stage is reached which corresponds to the sound velocity at the
nozzle outlet. In normal condensing turbines, in which the steam
flow is less than in top-turbine practice or in top turbines of moderate
size, the design is arranged so that the sound velocity is
exceeded under light load conditions.
With reference to Mr. Kroon’s comments concerning the phase
relation of the blade motion with respect to the time of application
of the steam load, the mathematical supplement shows that
for maximum amplification the steam load is applied at a point
where the blade is at zero displacement and starting on the backswing,
under conditions where the blade is loaded in an odd number
of half cycles. At this time of application the velocity is a
maximum in the backward direction. This is not the same as the
case of a suddenly applied load. With a high natural frequency
of the impulse-blade system, the time required to establish the
mean driving load on a blade should always be greater than a
half cycle of blade motion. In such cases, the maximum amplification
will take place when the mean driving load is established
in an odd number of half cycles with the phase relation here mentioned.
If, however, the mean driving load is established in less
than a half cycle, then the phase angle changes, depending on
the exact time relation, until in the limit case, in which the load
is applied in zero time, the maximum displacement of the blade
will be produced when the initial velocity is zero at the instant
the load is applied, as stated by Mr. Kroon. This limit case cannot
be realized in practice.
It is desirable to review the conditions that define primary
resonance. This summary was given in the abstract of the
paper as presented in Milwaukee and also in the mathematical
supplement by Mr. Emmert.
(а) The start of the application of the steam load takes place
at a point where the blade is at zero displacement and starting
on the backswing at which time the vibrational velocity is a
maximum backward.
(б) The full steam driving load is assumed to be reached in an
odd number of half cycles of blade motion. If the frequency of
the blade is such that there is not an odd number of half cycles
in the actual time of load application, the time is adjusted
slightly for purposes of calculation.
(c) The full steam driving load is assumed to be applied for an
even number of cycles of blade motion. The time adopted for
calculation purposes is adjusted slightly if necessary.
(d) The time at which the release of the load starts is assumed
to be at a point where the blade is at zero displacement with respect
to the static deflection and when the blade is starting on the
forward swing. In other words, the blade is moving at its maximum
velocity forward.
(e) The removal of load is assumed to be completed in an odd
number of half cycles.
(j) The idle portion of the arc is assumed to include an even
number of cycles.
With reference to Mr. Warren’s discussion, he stresses the
statistical study of impulse-blade performance from which his
company’s design practice has been evolved. As indicated in
the author’s comments on Mr. Kroon’s discussion, this same type
of statistical analysis has been a necessary standard practice for
many years with turbine-blade designers.
Referring to Mr. Warren’s comments on the continuous joining
of low-pressure blade-shroud bands, this has been the practice
of the Allis-Chalmers Company for many years. What the author
considers an exceptional record of reliable operation of low-pressure
blading justifies this construction. The very few cases where
last-row-blade trouble has occurred have been explained by wellrecognized
causes. In Allis-Chalmers practice, the low-pressure
blades are first shrouded in groups leaving, say, ten or more gaps
open between sections. The spindle disk, or in 3600-rpm machines
the entire spindle, is heated in a suitable oven to about
350 F. The body of the spindle is then protected with insulating
material and the blading allowed to cool for a predetermined
period, after which the open junctions in the shrouding are joined
by welding. When the temperatures in the spindle and the
blading are equalized, the shrouding is then in compression.
Experience has shown that there is little trouble in the breaking of
these joints. Where broken joints between shroud sections have
occurred, these have been traced to some recognizable cause,
such as the close proximity of ribs in the exhaust chamber, or
the unequal draft of steam through a bleeder connection in which
the extraction belt was not well proportioned.
Mr. Kimball'points out the dangerous effect of resonance of the
type classed by the author as secondary resonance. As indicated
by Mr. Kimball, resonance between the natural frequency of
the blade structure and the periodic force set up by the nozzle
flow can produce amplification factors of a high order. In fact,
if all terms of Equation [19] are considered for the case reviewed,
the amplification factor for true resonance is of the order of 1000.
Mr. Kimball further points out that the intensity of what he
terms “impulse excitation of bucket vibration” is responsible for
an intensity far below that arising from resonance produced by
even a moderate periodic force. This is extremely important, and
fortunately a condition that is capable of control by proper design.
The author is in full agreement both with Mr. Kimball and Mr.
Kroon regarding the desirability of experimental investigation of
the vibration of blading, although, in the author’s opinion, the
real value of such experimentation is not realized unless guided
by a sound background of fundamentals.
The author wishes to point out that the steam load of 350 lb
is applied at the mean diameter, not at the tip of the blade as
assumed by Mr. Kimball.
The cubical contents of the port section of the Northwest blade
are over three times the 0.6 cu in. used in Mr. Kimball’s discussion.
With reference to Mr. Rathbone’s remarks, impulse blading
must be designed to withstand primary resonance. Secondary
resonance can be avoided by design.
The author is in agreement concerning the important effect of
looseness of the blade roots on frequency and the difficulty of predicting
the frequency of a blading system in which looseness of the
blade roots enters the problem. In past practice, some cases of
partial-admission impulse-blade failures have been attributed
to undue looseness of the roots.
Where looseness of the roots of partial-admission blades is a
cause of failure, it is the author’s belief that the blades are not
correctly proportioned.
The difficulties associated with securing tightness of the roots
in their grooves under conditions of full speed and full operating
temperature are stressed in the paper, and the necessity of adopting
proportions for the blades and roots that will maintain stability
by the effect of centrifugal force is emphasized. If, for example,
the stress-amplification factor on the bending moment
is found to be 10, the centrifugal stability factor, expressed
as the ratio of centrifugal stabilizing moment to the static
value of the steam overturning moment, should be at least 15,
if stability is to be maintained.
Mr. Rathbone’s discussion gives this opportunity to stress

710 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
further the importance of maintaining design proportions so that
the rocking or chattering of the blades in their grooves due to
looseness can never occur.
With reference to the resonant conditions described by Mr.
Rathbone in connection with the last-row blades, the author is
in agreement with the general principles indicated in Fig. 35.
The periodic forces against which the blading must be protected
are probably not as much due to the variation in the steam driving
force from blade to blade, as to the effects produced when a given
blade passes through different pressure zones. The cylinder
design is planned to reduce to a minimum such pressure differences.
The effect produced when the low-pressure blades move
in and out of zones of different pressures is similar to that which
would be induced by torsional oscillation of the spindle.
The limiting frequency ratio above which trouble from blade
resonance is thought not to occur is in fair agreement with field
experience, which indicates that resonant vibration does not
cause trouble if there is close agreement with higher multiples of
the rotational frequency than 4 or 5.
In closing, the author wishes to express his appreciation to
the A.S.M.E. for the privilege of presenting this paper and to the
discussers for their valuable contributions.

Discussion of A tta c k on Steel in H igh-
C apacity Boilers as a R esult of O v e r­
heating D ue to Steam B lanketing
This paper was presented at the Spring M eeting in New
Orleans, La., February 23-25, 1939. Due to the fact that
preprints were not available, discussion at the tim e of presentation
was lim ited. Subsequent to publication in the
October, 1939, issue o f the Transactions it was decided to
continue discussion at the Spring M eeting in Worcester,
Mass., May 1-3, 1940. This discussion, together with the
authors' closure, appears herewith.
B y E. P. PARTRIDGE a n d R. E. HALL
P. B. P lace.1 Although the authors group the given examples
of tube failure under a general classification of steam
blanketing, there is sufficient difference in the nature of the
failures so that they may be arranged in at least two groups, i.e.,
failures caused or accompanied by cracking, and failures by corrosion
without cracking.
The first type of failure is characterized by cracking of the
metal in areas that are alternately wet and dry. This type of
failure is a form of stress corrosion in which attack is greatly accelerated
by alternate heating and cooling and occurs along the
bottom of the tube where the water laps the overheated metal.
A typical case occurred in a circulator tube between an upper
side-wall header and the steam drum. Cracking developed into
failure at the bottom of a tube bend which was above the normal
water level and normally out of the gas stream. Leakage of
gases through a superheater seal near the point of failure caused
local overheating, and a surging type of circulation, typical in this
part of a boiler, periodically quenched the overheated section.
The second type of failure is characterized by corrosion and
loss of metal. Such corroded surfaces have been found along the
tops of sloping tubes in the form of a narrow groove and also as
deep pits in the bottom of sloping tubes under sediment deposits.
Often the corroded surface is covered with soft smutty oxide
mixed in some cases with colored chemical deposits. This type
of failure appears to be due chiefly to direct chemical attack by
steam or water, or by salts which concentrate or deposit locally.
In the analysis of these failures we can perhaps start from the
following accepted facts:
1 The failures are not accompanied by a deformation of the
tube, except as such deformation may have resulted from excessive
heating caused by the insulating effect of deposited corrosion
products.
2 The corrosion product is iron oxide which requires a source
of oxygen.
3 The protective oxide coating, which normally covers a tube
surface, must first be broken down to expose fresh metal before
corrosion can develop.
The absence of deformation in a tube, the wall of which is corroded
to one half its original thickness over a long period of time,
definitely indicates that excessively high temperatures are not a
prerequisite of the corrosion. Temperatures in excess of 1200 F
are reported, which are believed either to be incidental or a
result of accumulation of corrosion products rather than a cause.
There has been built up in the last few years a popular belief
that very high metal temperatures are involved in these types of
failure. Dissociation of steam has been given in many cases as a
direct cause of corrosion and a source of oxygen. It is the writer’s
belief that dissociation has erroneously been construed to mean a
decomposition of water vapor by heat into oxygen and hydrogen
with resulting attack on the metal by the liberated oxygen. At
temperatures that will not cause deformation of a tube, the
amount of dissociation of steam by heat is very small. The oxygen
needed for corrosion is not oxygen freed by heat dissociation
of water vapor but is extracted from steam, water or oxygen-
containing salts by a reduction process. The metal is available
for that reduction only after its normal protective oxide coating
has been broken down.
The concept of temperature difference in steam-blanketed
tubes as given in Fig. 1 of the paper makes no allowance for
heat conductance around the tube wall. Any appreciable increase
in the temperature difference between the top and bottom
of a steam-blanketed tube must result in an increasing proportion
of the absorbed heat being conducted through the tube wall
around the steam space.
Thermocouple measurements of tube temperatures and metallurgical
analyses are given as evidence of high-temperature
conditions in conjunction with these failures. Chromel-alumel
couples peened into boiler tubes in gas temperatures above 1500 F
are quite dependent upon the method of installation for their
reliability. The failure of such couples is not due to overheating
where they are attached to the tube but due to oxidation and deterioration
of the lead wires exposed to the gases. During one investigation
of tube failures, we installed platinum thermocouples
adjacent to chromel couples on all of the boiler tubes in question.
At no time did any of the platinum couples register excessive
temperatures although many of the chromel couples did and we
found that the maximum period of reliability for chromel couples
in gas at 1700 F was 2 days. Though many apparently lasted
for weeks, any variations in temperature registered after 2 or 3
days were open to suspicion. Relative to metallurgical analyses,
it is common for reports to state that, because of changes in
grain structure, it is concluded that the metal has been at temperatures
in excess of 1000 to 1200 F for so many hours. In many
cases, because of corrosion, this may be quite true. However,
in one case reported in this manner, the sample had been taken
adjacent to a header located outside of the furnace where such
temperatures could not have existed.
The breaking down of the protective coating is slow and difficult
in an atmosphere of steam alone. The experience with
superheaters, experimental corrosion tests, and the commercial
production of hydrogen, all indicate that the action of steam on
hot metal is to form the protective coating and that the oxidation
slows up as the time of contact increases.
Water quenching,
chemical attack of salt deposits, or mechanical effect of rapid temperature
changes are necessary to break down the coating and
1 Test and Research D epartm ent, Combustion Engineering Company,
Inc., New York, N. Y. Mem. A.S.M.E.
allow corrosion to proceed.
711

712 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
In dry areas, formed by steam blanketing along the top of
sloping tubes, there is little quenching except by water splashed
against the top of the tube. However, the evaporation residues
from the splashing remain on the tube forming a concentrated
salt mat on the dry area. It is believed that the corrosion along
the top of screen tubes is chiefly by chemical attack. This attack
can be accelerated if the steam pocket periodically vents itself
and quenching stresses are added.
A similar concentration of chemical may develop under deposits
of loose sludge on steam-generating surfaces with resulting
chemical attack and corrosion under the deposit. Boiler water
may diffuse into such deposits at a rate sufficient to keep the
metal from burning and blistering but not at a rate sufficient to
wash away the evaporation residues. Once corrosion is started
by the chemical attack, the accumulation of products of corrosion
accelerates the action and, unless the tube is cleaned out,
the deposit may eventually insulate the metal sufficiently to cause
severe overheating and blistering. A failure of this type developed
in the bottom row of a sectional-header, high-pressure
boiler which was initiated by an excessive phosphate-sludge
condition in the boiler.
Periodic quenching of overheated areas prevents the accumulation
of chemical deposits and the likelihood of corrosion failure.
The quenching, however, sets up very severe local stresses because
of the initial rapid absorption of latent heat and the protective
coating is continuously cracked, exposing fresh metal
to attack. Tube failures under such conditions are crack
failures.
The expansion and contraction stresses in a thick-walled highpressure
tube, set up by overheating only 200 to 250 deg and
repeatedly cooling by the rapid evaporation of quenching water,
are very destructive to the protective oxide film. The metal
exposed by cracking of the film reduces the steam by chemical
action with liberation of hydrogen.
In conclusion it is believed that the presence of water and/or
its associated salts is necessary to the tube failures of the type
under discussion and that such failures can be initiated at tube
temperatures lower than 800 F. Improvement in boiler design,
and water conditioning; elimination of circulating sludge;
better distribution of heat absorption; and standardization of
conservative heat rates are the indicated trends toward the
control and elimination of failures of this type.
H. K k e i s i n g e b .2 The title of the paper implies that the type
of corrosion discussed in the paper came into prominence with the
use of high steam pressure. Many of the tube failures were accompanied
or caused by cracking of the internal surface of the
tubes at a point where the internal surface was alternately dry and
wet. The stresses produced by repeated quick-cooling result in
cracking. The cracks in the stressed metal provide a large and
clean surface for the action of boiler water, and the surface
corrodes rapidly. The question arises whether the very thick
walls of the tubes are not a major contributing factor in the
failures The tubes are selected for a given pressure, with a
factor of safety of about 10, whereas the steam drums are designed
with a factor of safety of 4.5. Why is the factor of safety
for the tubes so high compared with the factor of safety of the
drums Surely, if the drum failed, more damage would be
done than by failure of a tube.
Mathematical analysis of the temperature stresses under
normal conditions of operation indicate that the thick-walled
tube is subjected to severe temperature stresses which would
make failure by cracking almost inevitable, if the stresses were
s Engineer in Charge of Research and Development, Combustion
Engineering Company, Inc., New York, N. Y. Mem. A.S.M.E.
not moderated by the flow of metal.3 Steam drums are shielded
from heat because the shell is too thick and might be damaged by
heating. At the same time, thick-walled tubes are located where
they are subjected to severe heating. When inquiry is made concerning
the reasons for the very thick walls of tubes with a factor
of safety of about 10, the reason given is that the thickness is
somewhat irregular and that the thinnest part of the tube wall
after a certain amount of corrosion has taken place must still be
strong enough to withstand the pressure. Some of the tubes
have been corroded to less than one half their original thickness
without any apparent distortion of the corroded area due to
pressure. It appears that tubes with much thinner walls would
hold the pressure and be less subject to the type of corrosion
which is caused or accompanied by cracking of the inner surface
of the tube.
The engineer might benefit by the use of thinner-walled boiler
tubes. They would cost less and they might reduce appreciably
the failures from corrosion. He might also select tubes of
smaller diameter and thinner walls.
It is true that there have been some tube failures in lowpressure
boilers with thin-walled tubes, but it is probable that
the failures would have occurred in much shorter time had
thick-walled tubes been used. The failure is due to temperature
stresses and the stresses are greater in the thick-walled tube
than in a thin-walled tube. Lower stresses must be repeated
over a much longer period of time than high stresses before they
result in failure. A thin-walled tube may last years, whereas,
one with a thick wall may fail in a few weeks.
R. T. H a n l o n .4 Further tube failures were experienced in
the top tubes of the 4-tube-high headers. Thermocouple temperatures
at the top of these tubes showed a normal temperature
of from 600 to 620 F when the saturation temperature corresponding
to pressure was approximately 580 F.
When the boiler was operated under conditions of full rating
and pressure, the tube temperatures would suddenly increase to
approximately 1040 F at the top of the tube. This was probably
considerably lower than that at the bottom of the tube
where the damage occurred.
Reduction of drum pressure from 1300 to 1050 psi reduced
tube temperatures in all cases to normal values.
It was also observed that elevated tube temperatures increased
in frequency, elevation, and duration as the concentration
of total dissolved solids in the boiler water was increased.
Various arrangements of the drum baffles and steam washer
have been tried. The most effective baffle arrangement found
to date takes advantage of the velocity of the water in the drum
flowing from the uptake to the downtake tubes. The baffle is
so located, as shown in Fig. 1 of this discussion, that it
directs the flow of water to the sectional-header downtakes.
The effect of these baffle installations in combination with the
installation of restrictor plates upon the upper half of the lower
ends of the sectional-header tubes has been to permit the maintenance
of approximately 500 ppm total soluble solids in the boiler
water. It has been possible to maintain this concentration at
full boiler pressure and rating without experiencing elevated
tube temperatures or serious carry-over. It had not been possible
to maintain concentrations in excess of 100 to 150 ppm under
normal operating conditions previous to these installations.
During all of this period the feedwater to these boilers consisted
of turbine condensate, the make-up being evaporated in
low-pressure boilers. No further failures were experienced in thq,
upper sectional-header tubes.
3 “ Stresses in Boiler Tubes Subjected to High Rates of H eat Absorption,”
by W. L. de Baufre, Trans. A.S.M .E., vol. 55, 1933, FSP-
55-6, pp. 73-98.
4 Consolidated Edison Co. of New York, Inc., New York, N. Y.

DISCUSSION OF ATTACK ON STEEL IN HIGH-CAPACITY BOILERS 713
During October, 1939, after two months of operation with
externally treated city-water make-up, drifts of loose sludge
were observed near the downtake end of the bottom tubes of the
sectional-header bank. Subsequently, several tubes failed in
operation in this location, as is indicated in Fig. 2, herewith.
In every case of tube failure, large quantities of magnetic iron
oxide were observed at or near the point of failure. The appearance
and composition of a typical case are given in Fig. 3. Little
if any indication of blistering or swelling of the outside of the
tubes was observed. The apparent cause of failure was accelerated
corrosion on the inside of the tubes at points where deposition
had occurred. This deposition apparently caused localized
tube overheating due to insufficient heat transfer.
Previous to the time when these failures occurred, continuous
blowdown from the drum was employed for control of total
solids. Under ordinary conditions of operation, blowdown from
the drum is clear, indicating that, while soluble materials are
removed from the boiler, insoluble compounds are not. In order
to remove sludge from the boiler a procedure was, therefore, instituted
by which each boiler was taken out of service every 4
days and blown from the lower waterwall headers and the mud
drum.
The average alkalinity of the boiler water was also increased
from 15-25 to 20-30 ppm NaOH. This appears to have prevented
the further formation of a tough, adherent silica scale
upon the tube surfaces.
The amount of calcium-and-magnesium deposit was apparently
decreased by shortening the cycle of the external-treatment units.
In this manner the total quantity of scale-forming salts introduced
into the boilers was decreased and resulted in cleaner tube
surfaces.
The over-all effect of the increased boiler blowdown, higher
boiler-water alkalinity and the shortening of the cycle of the
primary treatment has been to arrest corrosion of the sectionalheader
tubes due to deposition of sludge and scale. These
procedures, however, have resulted in somewhat higher heat
losses and carry-over to the turbine, necessitating washing of the
blading.
W. L. W e b b .6 Inasmuch as the authors’ description of the
Logan ash-screen-corrosion experience was carried only to the
end of 1938, experience since that time may be of interest. As
the authors indicated, the ash-screen tubes on the west half of the
boiler were replaced during the November, 1938, shutdown
with 3-in-outside-diam 12.5-deg-sIope tubes of the same design
as those installed on the east half of the boiler in May, 1938.
Visual examinations of representative screen tubes in February,
1939, showed the usual brownish black areas along the tops
of the tubes, some being continuous narrow bands and others
discontinuous bands of varying width. Those in the tubes on
the west half of the boiler were more clear-cut than those in the
east, due possibly to thinner deposits of sludge. Day loads on
the boiler during the 3 months prior to this shutdown ranged between
800,000 and 1,000,000 lb of steam per hr.
Although no indication of active corrosion was found in the
screen tubes, these markings showed that corrosion might not
have been completely stopped. In order to produce more intimate
contact of the water with the top surfaces of the tubes,
twisted sheet-steel spirals were placed in all screen tubes during
the March, 1939, shutdown. These were constructed of pieces of
No. 18-gage steel, 5.5 ft long, having a width slightly less than the
inside diameter of the tubes and bent to give a complete 360-deg
twist in 1.5 ft of spiral. In addition to installing the spirals,
the lower burners were tipped upward 2.5 deg, giving a total
angle of 4 deg from the horizontal, in order to reduce further
the flame impingement on the screen tubes.
No major inspections of ash screen tubes were again made
until February, 1940. At that time, spirals were removed from
seven representative tubes. All previously affected areas were
1 American Gaa and Electric Service Corporation, New York,
N. Y.

714 TRANSACTIONS OF TH E A.S.M.E. NOVEMBER, 1940
evenly covered with a thin reddish colored sludge, no black oxide
whatever being evident. The calipering of these tubes indicated
no measurable metal loss. Except for a light coating of red
oxide, the spirals were clean and smooth and showed no indication
of corrosion.
As reported in the paper, hydrogen-evolution measurements in
saturated steam were started in March, 1938. The gases were
separated from the condensed sample by means of a separator
of the boiling-chamber type, similar to that used by Potter,
Solberg, and Hawkins in connection with their study of superheater-tube
metal and were analyzed with an Orsat. Hydrogen
was determined by the heated copper-oxide method. Usually
two 2-hr samples were obtained each day. Hydrogen values obtained
with this equipment appeared to have no direct relation
to corrosion rate. Tests of the gas-removal efficiency of this
separator, measured in terms of oxygen removal, showed it to
have too poor efficiency to be suitable for measuring accurately
the concentrations of hydrogen expected.
An automatically controlled separator for continuous operation,
patterned after a tray-type deaerator, was therefore designed
and constructed. After completion of preliminary tests,
this unit was placed in operation in April, 1939, and has been in
almost continuous service since then. Two gas samples are
collected per day over 8- and 16-hr periods, respectively. Unfortunately,
this new separator was developed too late to be used
during the period of severe screen-tube corrosion.
Since hydrogen-evolution measurements are being more extensively
made on steam samples from high-pressure boilers,
actual values in the Logan saturated steam may be of interest.
During the last year the evolution values have had a direct relationship
to load and are approximately 18 and 22 1 per 1,000,000
lb of steam at half load and full load, respectively. This relationship
has been substantially the same when feeding ferrous
hydroxide or sodium sulphite in proportion to dissolved oxygen
in the feedwater as a scavenger, or when no scavenger was fed.
No tests have been made to determine the effect of excess ferrous
hydroxide on hydrogen evolution.
Equipment is being provided at both Windsor and Twin
Branch for hydrogen-evolution measurements, the latter being
equipped with a two-point recorder. Values of hydrogen in the
feedwater, sampled at the economizer discharge (in one case
equal to one half to three quarters that in saturated steam), indicate
the necessity of sampling at both of these points, preferably
simultaneously, to show the quantity of hydrogen actually
evolved at the steam-generating surfaces.
The Logan boiler water pH is still being maintained between
9.6 and 9.8, since, under present load conditions, it is not considered
advisable to conduct experiments at higher pH values
and because evaporator carry-over and condenser leakage are at a
minimum. It is hoped that the pH ultimately can be carried at
about 10.8.
As was indicated by the authors, during the Logan screen-tube
corrosion experience, soluble boiler-water chemicals “hid out”
during high loads and were redissolved at low loads or when the
boiler was taken off the line. At one time when the phosphate
as P 0 4 in the boiler water was being maintained at 30 to 50
ppm, a total of 350 ppm of P 0 4 “hid out.” Since the tube corrosion
has been stopped, hiding out has ceased.
Experience with the Logan boiler as well as others can lead to
no other conclusion than that hiding out is an unfavorable sign,
indicating tube starvation or steam blanketing which may be a
forerunner of tube corrosion resulting from the reaction between
steam or water and the hot metal.
Uncompleted experiments with a 250-lb gage controlled-circulation
test boiler, having a single horizontal tube in a gas-fired
furnace, indicate that, with heat-absorption rates up to 125,000
Btu per sq ft of circumferential tube area, and with boiler-water
pH values of 9.6 to 12, the hydrogen evolution and top-tube
temperatures remain normal as long as the steam-water velocity
exceeds a critical value. However, as the velocity is decreased
below this critical value, the tube-wall temperature rises rapidly.
With increasing tube-wall temperatures, hydrogen evolution
increases, a higher hydrogen evolution occurring at a boiler
water pH of 11 than at a pH of 9.6.
The writer wishes to commend the authors for presenting the
case histories and particularly for the accuracy and completeness
of the Logan experience. It is evident that too few 1300-lb
boilers have escaped the type of tube corrosion which is known
not to have been caused by dissolved oxygen in the feedwater and
which is apparently beyond correction by chemical control of
boiler water. A better understanding of this type of corrosion,
both by the boiler manufacturer and the purchaser, should lead to
a definite solution of this problem.
L. E. H a n k i s o n .* The discussion of any paper presented over
a year ago should offer a very excellent opportunity to substantiate
or disprove the various observations, theories, and conclusions
presented in the paper; and in so far as our experience
goes, the findings and theories advanced by the authors have
been borne out to an extraordinary degree. Therefore, the
writer believes the paper under discussion should be considered
as authoritative on the various ideas brought forth.
With regard to the case history reported for Springdale, the observations
and causes of the troubles outlined in the paper have
been confirmed, and the loss of steam-generating tubes stopped,
we believe, by the removal of the cores that were placed in
the tubes to promote circulation, and the installation of distributing
apparatus in the headers to prevent steam binding.
The manufacturer of the steam-generating equipment furnished
to Springdale is to be highly commended on the handling of the
various problems, the analysis of troubles as they developed,
and the application of the proper corrective measures for producing
the desired results. As the coal conditions and boiler-water
concentrations with which we are now operating are considered
much more severe than existed at the time of our troubles, we
feel justified in the belief that our present satisfactory conditions
will continue.
We have always held to the theory that the steam generator
should be so designed as to allow the use of properly treated feedwater,
and that undertreating of feedwater in an effort to remedy
the shortcomings of boiler design should not be tolerated. The
paper suggests that, by the use of inadequate treatment of the
boiler water, the detrimental effects of steam binding and the
wasting away of boiler metal may be retarded in the dry areas,
but that lower alkalinity may possibly cause deterioration
throughout the boiler. It seems that metal cracking is likely to
take place under these conditions, and that, instead of correcting
the trouble, the time is only prolonged until general deterioration
takes place. On the theory that proper boiler-water conditions
should be maintained, we carried sufficient alkalinity and
conditioning chemicals properly to treat the high-pressure boilers
at Springdale, even while undergoing tube troubles. Our experience
shows that correction of the circulation and steam-binding
conditions is all that is required to permit satisfactory operation
with properly treated boiler water.
We have had further opportunity at Springdale to confirm the
conclusions outlined in the authors’ paper. Because of unexpected
characteristics of the coal, it was considered advisable to
make the walls of the secondary furnace entirely water-cooled by
adding extra tubes to the back and side walls. I t was thought at
6 Efficiency Engineer, W est Penn Power Company, Pittsburgh, Pa.
Mem. A.S.M .E.

DISCUSSION OF ATTACK ON STEEL IN HIGH-CAPACITY BOILERS 715
the time that the water-supply and steam-relieving conductors
already installed were adequate to take care of this extra waterwall
surface. After 17 days of operation of the full waterwalls,
tube failures were experienced at each end of the back wall, the
tubes being the fourth from their respective ends. The ruptured
sections were replaced and, after 18 days, one of the replaced
tubes failed. Also, one of the side-wail tubes failed at
the same time, this tube also being the fourth tube from the
cooler end of its header. At a later period, the same tube in the
opposite side wall failed, all of which seems to bear out the hypothesis
that these failures were caused either by inadequate
water supply to the waterwalls or by insufficient steam-relieving
tubes from their upper headers.
In an effort to overcome these troubles, the water-supply and
F ig . 5
C ro ss S e c t io n o f T u b e a t R u p t u r e M a r k e d t o S h o w
L o c a t io n o f M ic r o g r a p h s
F ig . 7 M i c r o g r a p h a t P o i n t 4 (F ig . 5 ), S h o w in g L o c a t i o n o f
M e t a l S t r u c t u r e C h a n g e
(This point ia where oxidation and thinning begins to show on inside
of tube; X 100.)

716 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
steam-relieving tubes of the headers were increased by approximately
100 per cent. Since these additions, there has been but
one failure and this was attributed to oxidation that occurred before
the extra water-supply and steam-relieving tubes were installed.
These failures were in the nature of metal cracking as
described in the authors’ paper. As all the tubes were in a vertical
position, the conclusion is reached that vertical waterwalls,
having inadequate supply and delivery connections, can have
certain tubes, presumably the ones in the hottest zone, delivering
an upward circulation; others, in the cooler portions of the wall,
a downward circulation; while some tubes probably have little or
no circulation.
At the time of these failures we were carrying concentrations
in the boiler water of 1000 ppm total solids. The standard is
now set at 2000 ppm and it is anticipated that this limit will be
increased in the near future. The boiler water carries excess
phosphate and hydroxide, and from 3 to 10 parts of sodium sulphite.
Some of this highly concentrated boiler water is recirculated
to the economizer inlet so that the economizer or first
heat-absorbing surface of the unit is at all times protected against
oxygen corrosion by the alkalinity and sulphite. Boiler feedwater
having a pH of about 8 is raised to a pH of 9.5 or 10 by this
recirculation. As a constant quantity of boiler water is recirculated,
the pH of the water through the economizer naturally
varies with the rating on the boiler and the concentration of the
boiler water.
Regarding the decomposition of sodium sulphite, we have
definite proof of its stability under normal 1300-lb service.
During the time when steam binding was taking place, we were
able to predict the failure of boiler tubes by the disappearance of
sodium sulphite in the boiler under question. While the sulphite
content of the other boilers in the plant would remain practically
constant, the one suffering deterioration would change
the sulphite to sulphate, and the chemical balance of the boilerwater
constituents fully accounted for all the sulphite which had
formerly been present.
It has been found at Springdale that, by maintaining an excess
of sulphite, we are amply protected against the admission of
oxygen which seems to appear when adding to or taking off
equipment. Even though the Winkler test usually shows zero
oxygen, there is a constant slow disappearance of the sulphite
and this is particularly noticeable over the week end when severe
load changes are experienced.
Further confirming the observations made by the authors in
their paper, sections of the vertical waterwall tubes that failed
show metal cracking, and micrographs of specimens from the
overheated surface emphasize a definite zone where overheating
caused the metal structure to change. These pictures are very
interesting in showing that the metal structure undergoes a
drastic change at the line where thinning by oxidation begins.
Fig. 4 shows the ruptured part of a waterwall tube after 17
days of service. In Fig. 5, the cross section of this tube at the
rupture is marked to show the location of the micrographs in Figs.
6, 7, and 8. The micrograph in Fig. 6 exhibits the pearlitic structure
of the original tube metal at point 3. At point 4, where
oxidation and thinning begin to show on the inside of the tube,
the structure shows a change, as indicated in Fig. 7. Fig. 8
illustrates the changed metal structure which is typical of the
metal from point 5 to the rupture.
The authors refer to the “hiding out” of boiler-water chemicals
and show graphical records of this more or less detrimental factor
to boiler-water conditioning. While Springdale has not experienced
this “hide-out” in so far as we are aware, we are experiencing
it in other plants on the system. A search in one of our high-pressure
boilers for this hide-out disclosed a residue of boiler-water
solids in the intermediate uptake waterwall header 30 ft below the
Fio. 9 C h e m ic a l D e p o s its in I n t e r m e d i a t e U p t a k e W a t e r w a l l
H e a d e r , W h ic h I s L o c a t e d 30 F t B e l o w C e n t e r L i n e o f B o i l e r
D ru m
normal water level of the boiler. This deposit was found adhering
to the surfaces of the header and tubes when the boiler
was drained, more than 22 hr after it had stopped steaming.
The beautiful frost-like patterns of chemical deposit are shown in
Fig. 9.
A very important factor in the satisfactory operation of the
Springdale high-pressure boilers is the steam-purifying equipment
which was described in a paper by M. D. Baker at the
same meeting at which the authors’ paper was presented. This
equipment is continuing to give first-class performance and
shows no deterioration. The clean steam we are able to obtain
with high boiler-water concentrations is, in the greatest measure,
attributable to this purifying apparatus.
The highly concentrated boiler water, carrying its proportional
amount of sulphite, phosphate, and alkalinity that is
standard for Springdale, has produced the following beneficial
results:
1 Complete economizer protection, as provided by recirculating
boiler water through the economizer.
2 Clean metal surfaces as regards scale deposition and corrosion,
and reduction of the amount of deposit to a point where this
trouble is negligible.
3 A clean, crystalline type of sludge deposit.
4 Reduced blowdown, relieving the load on the evaporated
make-up and heat-exchanging equipment.
These desirable conditions are all obtained at a decrease in
cost of operation and maintenance, and a paper is being prepared
for presentation in the near future in which will be given
more detailed information on steam quality, economies gained
by higher boiler-water concentration, proof of stability of sodium
sulphite, and the benefits derived by recirculation of boiler-drum
water through the economizer.
A u t h o r s ’ C l o s u r e
The discussion certainly continues the spirit in which the
paper was written—that of a cooperative venture in which evidence
from many sources was fitted together to produce as rational
and consistent a picture as possible. In this same spirit,
the authors venture a brief correlation of the preceding comments.
Fundamentally, what has been under discussion is the reaction
between iron and water and the effect upon this reaction of dissolved
substances and of temperature. As Mr. Place has pointed
out, iron is a strong reducing agent. The chemical driving force
for iron to take oxygen from water and liberate hydrogen is very
great, even at room temperature, but the iron oxide, which is

DISCUSSION OF ATTACK ON STEEL IN HIGH-CAPACITY BOILERS 717
one product of the reaction, quite literally gets in the way and
slows down the reaction to such an extent that it is scarcely noticeable.
Anyone can observe the progress of this reaction, however,
by allowing a quantity of very finely divided metallic iron to
stand in distilled water in a small bottle on his desk. On shaking
this bottle once a day, small bubbles of gas will be observed to
rise from the powder. This will continue day after day; a test
after a sufficiently long time will demonstrate that the gas is
hydrogen.
This same oxidation of steel by water, or reduction of water by
steel, goes on continuously in every boiler, but it is only when the
oxide resulting from it fails to retard the reaction that the boiler
operator has a problem on his hands.
The protective oxide film can be at least partially destroyed or
rendered more permeable by very high concentrations of sodium
hydroxide. Because it is brittle and differs in thermal expansion
from steel, it may also be cracked by repeated temperature
changes. Anything which causes excessive local concentration
of a boiler water containing some caustic soda or repeated overheating
and quenching of a tube surface is thus potentially
dangerous. To the authors, steam blanketing seems inevitably
to tend to produce one or the other or both of these effects.
The factor of caustic attack has not been questioned in the
discussion, but Mr. Place has gently indicated his disbelief in the
validity of thermocouple measurements or of changes in the
microstructure of steel as indications of excessive temperature
in the top of a steam-blanketed tube. Base-metal thermocouples
admittedly do fail all too rapidly when exposed to the conditions
in a boiler furnace, yet recent studies at the Bureau of Standards
show a maximum error of only 21 F, and this low rather
than high, when chromel-alumel couples were heated for long
periods of time in air. This maximum error was found after a
couple had been exposed for 200 hr at 2200 F .T Failure occurred
before 300 hr.
While the authors attach more significance than does Mr.
Place to the temperature measurements mentioned in the paper,
they feel that the microstructure of the steel is a still more certain
criterion of overheating. This is demonstrated particularly
well by the photomicrographs presented by Mr. Hankison correlating
the change in structure in the tube wall with the zone of
damage along the internal surface. Perhaps the sample which
inspired Mr. Place’s lack of faith in metallographic evidence actually
had been overheated at some time in some unrecorded
manner.
Mr. Place has argued that, if a tube were seriously overheated,
it would deform to a greater extent than has been observed in
many cases where grooving along the ceiling produced failure.
In a tube of 3.5 in. outside diam with a wall thickness of 0.5 in.
subjected to an internal pressure of 1400 psi, the nominal stress
tending to rupture the tube is, however, only 3500 psi for metal
in the longitudinal section of the wall. That creep of a lowcarbon
steel subjected to this stress at a temperature of 1000 F
is slight has been shown by White, Clark, and Wilson.* They
observed that SAE 1015 steel held at 1000 F for 16,000 hr under a
load of 4000 psi showed an average rate of creep of only 0.043
per cent per 1000 hr.
It must be remembered also that a narrow band of steam
along the ceiling of a tube would produce only a narrow band of
overheated metal and a narrow groove, as in Fig. 15 of the paper.
7 “Stability of Base-Metal Thermocouples in Air From 800 Degrees
to 2200 Degrees Fahrenheit,” by A. I. Dahl, Research Paper
R P 1278, N ational Bureau of Standards, Journal of Research, vol. 24,
1940, pp. 205-224.
8 “ Influence of Time at 1000 F on the Characteristics of Carbon
Steel,” by A. E. W hite, C. L. Clark, and R. L. Wilson, Proceedings of
the American Society for Testing M aterials, vol. 36,1936, p art 2, pp.
139-156, see Fig. 1, p. 141.
Deformation in such a case might be limited to a strip not more
than 1 in. wide. Where the steam blankets more of the tube
ceiling, as in the cross section shown by Mr. Hankison, there,
usually is definite stretching of the tube wall prior to failure
resulting in longitudinal cracks in the internal and external layers
of oxide.
Overheating of a tube is, of course, not a new or unique occurrence.
While we have thought for years in terms of the overheating
due to deposits of solid on a heat-transfer surface, much
less consideration has been given to the insulating effect of
a blanket of steam, which was the starting point of the paper.
The experiences at Springdale and at Waterside described
in the paper and the discussion by Mr. Hanlon bring us back to
the fact that sludge settling out on the bottom of a tube can interfere
with the transfer of heat from metal to water just as seriously
as steam along its ceiling. The sludge which caused the
failures at Waterside, reported by Mr. Hanlon, like that which
led to the failures along the bottom of the cored tubes at Springdale
was not, however, the relatively light calcium phosphate
sludge which results from phosphate conditioning but, instead, a
heavy sludge of magnetic iron oxide and metallic copper. Unless
powdered metallic iron or ferrous hydroxide is being introduced
intentionally into the boiler feed, such a sludge can result only
from undesirable attack of water on steel in some region within
the boiler, not necessarily the region where the sludge is found.
The initial settling out of the sludge on the bottom of the tubes
described by Mr. Hanlon may have been favored by the presence
of restrictors at the inlet ends of the lower tubes, as it was apparently
by the cores in the tubes at Springdale. Overheating of
the portion of the tube covered by the sludge could then have
led to progressive oxidation of the tube wall.
Attempts to increase flow in one particular region of a boiler
by specifically retarding it in another region seem, in a number
of cases, to have transferred trouble rather than to have prevented
it. Of the various expedients which have been tried to
minimize steam blanketing in tubes through which there was
positive circulation, the spirals introduced into the screen tubes at
Logan, described by Mr. Webb, seem the most promising.
The suggestion of Mr. Kreisinger that tubes in high-pressure
boilers are too thick-walled for their own good merits further
study. Reduction in wall thickness will not, however, change
the heat input to a tube or minimize the formation of a steam
blanket within it. While cracking of tubes subjected to repeated
overheating and quenching may be slower with thin than
with thick walls, as suggested by Mr. Kreisinger, the paper records
extensive damage to the thin-walled tubes of the low-pressure
boilers in the Beacon Street Heating Plant in 8 years. The
authors believe that grooving along the top of a tube might occur
as rapidly in a thin-walled as in a thick-walled tube.
It was first demonstrated at Port Washington that grooving
of the ceilings of inclined tubes could be minimized without
mechanical changes by eliminating caustic alkalinity from the
boiler water. This expedient was adopted at Logan, but the
slope of the screen tubes affected was also increased, the spirals
were inserted to throw water against the tops of the tubes, and
the burners were tilted upward to reduce flame impingement.
At Springdale, no change was made in water conditioning, but
the difficulties first encountered were eliminated solely by mechanical
changes. An appropriate conclusion to the question of
the relative importance of the chemical and the mechanical factors
has been written into the record by the announcement that,
after operation for 2Vs years with no caustic alkalinity in the
boiler water, extensive damage of the original type was recently
discovered in the screen tubes at Port Washington.*
9 “P o rt W ashington Sustains Its Economy,” Combustion, vol. 11,
no. 7, 1940, pp. 33-34.

D eterm ination of the P u rity of Steam by
G ravim etric and S pectrographic M ethods
By M. C. SCHWARTZ,1 W. B. GURNEY,2 a n d T. E. CROSSAN3
This is the first paper in a series of three covering in ­
vestigations on the determ ination o f purity o f steam . In
it the authors discuss the m echanism o f steam contam ination
and m ethods of measuring steam purity classified
according to im purities present. Principally, however,
the subject m atter is devoted to the apparatus and procedure
followed in making gravimetric determ inations of
the residue from evaporated samples of condensed steam
with the results obtained. As a check on these analyses,
estim ates are made for solids in steam from turbine-blade
deposits. The results of spectrographic analyses are also
given.
THE purity of steam produced from a boiler is a function
of the composition of the boiler water and the boiler feedwater,
as well as the mechanical design of the boiler in
which it is produced. The production and consumption of huge
amounts of steam at high pressures and temperatures have demanded
the utmost in purity of the steam leaving the boiler, because
of the economic importance of maintaining clean superheaters,
turbines, and process equipment. Any changes made to
increase the purity of steam are thus quite important and, particularly,
since they may involve considerable expense, should be
undertaken with due consideration of the methods used to determine
the purity of the steam. Investigation of the methods
of determining steam purity has been made at the Louisiana
Station and has been applied in the control of steam production
at this station.
M e c h a n is m o r S t e a m C o n t a m in a t io n
Steam becomes contaminated with gases or moisture containing
dissolved or suspended solids present in the boiler water or in
the feedwater which may be used as a scrubbing medium for the
steam. The contamination of steam results from the continuously
occurring phenomenon of foaming or from the intermittent
and irregular phenomenon of priming. An excellent summary
of the factors contributing to foaming and priming has been presented
by Powell (l).1 Under boiler design this author covers such
considerations as steam disengaging space, steam storage space,
maximum water space, size of steam header, velocity of steam,
various obstructions in the steam drum, and boiler pressure.
1 Research Chemist, Gulf States Utilities Company, and Research
Associate in W ater Technology, D epartm ent of Chemical Engineering,
Louisiana State University.
2 Efficiency Engineer, Gulf States Utilities Company, B aton Rouge,
La. Mem. A.S.M.E.
3 Superintendent, Gulf States Utilities Company, B aton Rouge,
La. Mem. A.S.M.E.
4 Numbers in parentheses refer to the Bibliography a t the end of the
paper.
Contributed jointly by the Gulf States Utilities Company, Louisiana
Station, B aton Rouge, La., and the Engineering Experim ent
Station, Louisiana State University, and accepted by the Joint Research
Committee on Boiler Feedwater Studies and presented at
the Spring Meeting, Boiler Feedwater Sessions, of T h e A m e r i c a n
S o c i e t y o f M e c h a n i c a l E n g i n e e r s , W orcester, Mass., M ay 1-3,
1940.
N o t e :
Statem ents and opinions advanced in papers are to be
understood as individual expressions of their authors, and not those
of the Society.
Boiler operation covers such matters as rate of evaporation, load
swings, and water level. Water treatment covers such items as
composition and concentration of dissolved and suspended solids
in the boiler water and in the feedwater, as well as the chemical
condition at the metal surfaces exposed to the water.
M e t h o d s o f M e a s u r i n g S t e a m P u r it y
Methods of measuring steam purity may be classified according
to the impurities present. For example, if an appreciable
amount of moisture is being determined, the steam calorimeter
can be used. If all the soluble electrolytes, arising from the
solids or gases in the steam, are being determined in the condensate,
then the electrolytic-conductivity method can be used. If
specific chemical substances, such as ammonia, carbon dioxide,
sodium chloride, and the like, are being determined, then specific
chemical methods for these substances can be used. If the
total solids present in the steam is being determined, then the
residue after evaporation of the condensate can be weighed or,
alternatively, it is apparently possible to deposit the solids on a
surface and determine them in this manner. It is useful to consider
the subject from the viewpoint of the impurity, because it is
quite possible that some method, perhaps satisfactory in its own
way, serves no really useful purpose in determining why the deposits
in the equipment are forming and in what way they may be
avoided.
G r a v im e t r ic D e t e r m i n a t io n o f t h e R e s i d u e F r o m E v a p o ­
r a t e d S a m p l e s o f C o n d e n s e d S t e a m
The gravimetric determination of the total solids, including
thus soluble and insoluble substances, in the steam by weighing
the residue after evaporation of a sample of steam condensate
can probably be considered the standard procedure. The
minute weight of the residue offers no serious drawback since it
may be determined readily with a sensitive balance. The problem
of evaporating the sample of condensate free from subsequent
contamination offers more difficulty. Sufficiently slow evaporation
of the sample to avoid loss by spattering results in a serious
loss of time and it is this factor which so far has prevented this
method from becoming adaptable to routine use.
Although the preparation of pure water has been the subject of
numerous researches in the laboratory, the investigators very seldom
have determined the residue on evaporation. However,
Acree and Fawcett (2) evaporated 301 of pure water, contained in
a quartz flask, in a platinum dish. The maximum residue was
never more than 0.15 ppm and was as low as 0.03 to 0.06 ppm.
Ulmer (3), using condensed-steam samples from a commercial
boiler, determined the residue on evaporation to be 0.19 to 1.01
ppm, depending upon the boiler-water concentration.
A semimicro apparatus has been developed for evaporating 2-1
samples of condensed steam. Ulmer (3), Powell (4), and Seyb (5)
have considered the problem of evaporating condensed-steam
samples. Fig. 1 shows the manner in which the evaporator works.
The apparatus was built upon a triple-beam balance, of which
Sargent No. 3455 is an example. The knife-edge mechanism, brass
lever arm, and transite holder for the Pyrex micro bell jar were
built in the University shops. The Pyrex jar is Corning No. 6880.

720 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
The rubber tubing is pure gum. A 2-1 Pyrex Erlenmeyer flask,
protected from carbon dioxide with a soda-lime-sodium-hydroxide
guard tube, contained the sample of condensate. Pyrex tubing
was used to convey the water under constant head from the flask
to the platinum dish beneath the bell jar. Block-tin tubing was
available for use in place of Pyrex tubing if it seemed necessary,
but the results did not seem to warrant any change. A piece of
platinum wire was sealed to the Pyrex tubing so that hot water
would not come into contact with the Pyrex-glass tubing. A
200-w 2*/2-in. Chromalox heating unit was used and the heat input
was controlled by a rheostat. The platinum dish used,
weighing 10 g, was the lightest obtainable commercially for its
size. It was 2 in. in diam and 1 in. deep. This dish was supported
upon a porcelain ring. The side arm of the Pyrex bell jar was
connected to a Fisher air ejector. The trap in this line prevented
the possibility of any water being sucked back.
In operation, the lead weight on the balance beam is initially
adjusted to balance the weight of the platinum dish and water as
well as the Pyrex jar and other accessories. For an actual run,
control is made by using small weights on the brass lever arm.
A change of about 10 g starts or stops the flow of water. After
an initial adjustment of the rate of flow, the evaporation always
continues with practically no attention for the time necessary,
which is about 48 hr.
The total solids have been determined on condensed-steam
samples from the station boilers. Test data are presented in
Tables 1 to 3.
From the results of Tables 1, 2, and 3, it will be seen that for
this station the totals of solids in the condensate from saturatedand
superheated-steam samples are practically identical. The
copper degasifier, which will be described later and which removes
carbon dioxide from the samples, gives results approximating
those obtained on samples subjected to no treatment for degasification.
In all cases, the samples were collected within several feet
of the sampling points.
From previous experience with deposits forming on glass conductivity
cells, as well as by the deposits formed in the platinum
dish, it was obvious that either copper or iron or both were present
in the condensed steam. Tests for copper with diphenylthiocarbazone
or dithizone on the condensed steam, and on
samples concentrated by evaporation, gave negative tests.
However there seems to be no doubt that high steam velocities
will cause erosion of copper tubing. Whether or not these minute
particles of copper will dissolve seems to depend upon a number
of factors, one of the most important of which is the rate of flow
of condensed steam through the sampling system. Interesting
but old data on the solubility of copper in distillled water are
given by Carnelley (6).
Semimicro determination of the iron in the residues from
samples Nos. 9 and 10, Table 2, showed 0.23 and 0.25 ppm
Fe20», respectively, out of a total of 0.3 ppm solids. Since these
samples were collected in a most satisfactory manner, as far as
avoiding possible contamination of iron from the sampling line is
concerned, the conclusion seems logical that minute particles of
iron are in suspension in the steam and are so finely divided that
they show no tendency to settle out in the condensed steam.
These iron particles are not visible in the solution but are removed
by ultrafiltration. There is no reason to assume that
these iron particles will affect the electrolytic conductivity of
the solution of condensed steam. Upon the basis of these results
of gravimetric determination of the residue from evaporating
condensed-steam samples, coupled with determination of iron
and of copper in these residues, there is apparently no more
than about 0.1 ppm soluble solids in the steam at the Louisiana
Station.
E s t im a t io n o f S o l id s i n S t e a m F r o m T t jr b in e - B l a d e
D e p o s i t s
The values obtained for the purity of steam by any given
method should correspond to the actual practical data observed
in the plant. For example, the amount and kind of deposits observed
on the turbine blades should be in proportion to the indi-
T A B L E 4
T otal wash w ater leaving turbine, lb ................................................... 8181
G ravim etrically determ ined to tal solids in wash w ater, lb ............ 1.5786
T otal steam passed through turbine, lb .............................................. 340340000
Solids on turbine blades washed off, p p m .......................................... 0.005
T A B L E 5
T o tal wash w ater leaving turbine, lb ............................................... 9146
G ravim etrically determ ined to tal solids in wash w ater, lb ..........
T otal steam passed through turbine, lb ..............................................
8.34755
775483000
Solids on turbine blades washed off, p p m .......................................... 0.01
T A B LE 6
T otal wash w ater leaving turbine, lb . . . . ...........................................
G ravim etrically determ ined to ta l solids in water, lb ......................
4949
1.4882
T o tal steam passed through turbine, lb .............................................. 388975000
Solids on turbine blades washed off, p p m . ........................................ 0.004

SCHWARTZ, GURNEY, CROSSAN—DETERM INATION OF THE PURITY OF STEAM 721
cations obtained by any one of the usual methods of determining
steam purity. Data of this kind, as useful as they are, do not
furnish a continuously available record of steam purity as a guide
to efficient operation.
The authors have attempted to determine the solids present in
the superheated steam by washing a turbine with condensate,
while the machine was rotating at slow speed under no load.
Data are presented in Tables 4, 5, and 6 for three different turbines.
Although the solids were removed from the turbine by washing
with condensate, all of the deposit so obtained was not soluble.
Table 7 shows the relative amounts of insoluble and soluble material.
T A B L E 7
D ata of
Table 4
Table 5
Table 6
R E L A T IV E A M O U N TS O F IN SO L U B L E A N D SOLUBLE
M A T E R IA L
Solids, w ater
soluble, lb
1.5462
7.3534
1.2081
Solids, w ater insoluble,
lb
0.0324
0.9942
0.2801
T o tal solids,
lb
1.5786
8.3476
1.4882
S p e c t r o g r a p h ic A n a l y s is o f C o n d e n s e d S t e a m
The spectrographic analysis of condensed steam was made at
the spectrographic laboratory of the Massachusetts Institute of
Technology. The authors in order to avoid contamination of the
sample by the container, selected what was believed to be the
most inert substance besides platinum, which was prohibitive in
cost. The sample was collected in a 500-ml silver Erlenmeyer flask
with a silver stopper. A qualitative analysis of the sample showed
the presence of the following elements in diminishing amounts:
silver, copper, iron, magnesium, sodium, and silicon. In addition,
minute traces of boron, lead, and calcium were found.
Phosphorus was not detected even with extreme concentration of
the sample; however, this element is one to which the spectrographic
method is less sensitive. A quantitative analysis gave
the results shown in Table 8.
T A B L E 8 Q U A N T IT A T IV E A N A LY SIS O F C O N D E N SE D -ST E A M
SA M PL E
Substance
Ppm
C o p p er........................................................................ ... 2 .4
Iro n .............................................................................. ....0.06
S o d iu m .......................................................................... 0.14
M agnesium ............................................................... ....0.056
Silicon..............................................................................0.015
The presence of such large amounts of silver and copper was
entirely inconsistent with the gravimetric results, as well as with
the known electrolytic conductivity of this sample. Apparently
the silver flask was the source of this unusual contamination.
Accordingly another sample was collected for spectrographic
analysis, this time in a Pyrex glass-stoppered bottle. Table 9
shows the results obtained.
TA B LE 9 Q U A N T IT A T IV E A N A LY SIS O F A D D IT IO N A L C O N ­
D E N SE D -ST E A M SA M PL E
Substance
Ppm
C o p p er................................................................. 0.05
S ilic o n ................................................................. 0.10
C alcium ................................................................. less th a n 0.01
The drop in copper concentration from 2.4 to 0.05 ppm with
the change in the container shows contamination from the silver
flask, evidently from the silver solder used in fabrication. Correspondingly,
the silicon concentration changed from 0.015 to
0.1 ppm, evidently due to the effect of the glass container.
By combining the results of both analyses, we derive the probable
composition of the solids present in condensed steam, as
shown in Table 10.
On the basis of the results shown in Table 10, one would expect
a value for the total solids of condensed steam to be higher than
0.33 ppm, the exact value depending upon the actual negative
ions associated with the positive ions shown in the table. It
seems somewhat surprising that the results shown by spectro-
TA B LE 10
R ESU L TS O F C O M B IN IN G ANALYSES O F TABLES
8 A N D 9
Substance
Ppm
Sodium ......................................................................... 0.140
C alciu m ...................................................................... 0.010
M agnesium ................................................................ 0.060
Silicon........................................................................... 0.015
Iro n ............................................................................... 0.060
C o p p er........................................................................ 0.050
graphic analysis should be greater than those obtained by evaporation
and subsequent weighing, particularly since it has been
assumed that the evaporation might probably be affected by
contamination during the long period required. For further investigation
using the spectrographic method, it will be desirable
to establish the kind and amounts of impurities which may be
present in the electrodes which are used in the analysis. That
this type of error may exist is quite probable because of the fact
that both calcium and magnesium are absent from turbine-blade
deposits. In addition, it is well known that the electrodes used
in spectrographic analysis are freed from traces of impurities only
with difficulty. Traces of boron, silicon, copper, calcium, magnesium,
and sodium usually exist even in treated electrodes (7, 8).
Hilger H.S. carbon rods of very high purity contain traces of
boron, calcium, copper, magnesium, and iron (9).
The sensitivity of the spectrographic method, as is evident
from the foregoing analyses, makes it an extremely valuable
tool with which to attack the problem of steam purity. From
data presented by Owens (10) the authors have calculated the
sensitivity of the spectrographic method to be as shown in
Table 11.
T A B LE 11
S E N S IT IV IT Y OF S P E C T R O G R A P H IC M E T H O D
Sensitivity
M inimum concentration
(inorganic liquid)
(using 0.1 ml), ppm
A1.................................................. 0.25
C a ............................................... 0.2 0
M g................................................ 0.15
S i .................................................. 2.50
C r.................................................. 0.1 0
C u ................................................
F e ............................................. .. .
0.0 5
0.05
M n ................................................ 0.01
N a ................................................ 0 .3 8
P b .................................................. 0 .1 0
S r .................................................. 0.05
Even with the sensitivity shown in Table 11, some concentration
of the condensed-steam samples is required in order to obtain
indication of some of the substances present in the sample.
In conclusion it should be remembered that spectrographic
analysis will give the sum of the soluble as well as insoluble solids
present in the condensed steam.
A c k n o w l e d g m e n t
The authors wish to thank H. C. Leonard, vice-president of the
Gulf States Utilities Company, and Dean L. J. Lassalle, director
of the Engineering Experiment Station, Louisiana State University,
for permission to publish this material. Grateful acknowledgment
is made to E. B. Powell, consulting engineer of
Stone and Webster Engineering Corporation, and to S. T. Powell,
consulting chemical engineer, for their advice, highly valued suggestions,
and encouragement throughout the course of the work
being reported. Members of the chemical staff of the Gulf
States Utilities Company assisting in this research were L. Young
and J. C. Hill.

722 TRANSACTIONS OF TH E A.S.M.E. NOVEMBER, 1940
B IBLIOGRAPHY
1 “ Steam Contam ination,” by S. T . Powell, Combustion, vol. 9,
no. 3, September, 1937, pp. 36-40.
2 “ The Problem of Dilution in Colorimetric H -Ion Measurem
ents,” by S. F. Acree and E. H. Faw cett, Industrial and Engineering
Chemistry, analytical edition, vol. 2, 1930, pp. 78-85.
3 “ Determ ination by the E vaporation M ethod of Small Amounts
of Dissolved Solids in W ater Such as Condensed Steam From
Boilers,” by R . C. Ulmer, Proceedings of the American Society for
Testing M aterials, vol. 39, 1939, pp. 1221-1232.
4 “ Design and Development of A pparatus for M easurement of
Steam Quality by Electrical C onductivity M ethods,” by S. T. Powell,
H. E. Bacon, Jr., I. G. McChesney, and F. Henry, Transactions of
the American In stitu te of Chemical Engineers, vol. 33, 1937, pp.
116-138.
5 “ Analysis of Condensates," by E. Seyb, “ Von W asser,” vol. 8,
p art 2, 1934, pp. 169-172.
6 “ On the Action of W ater and of Various Saline Solutions on
Copper,” by T. Carnelley, Journal of the Chemical Society of London,
vol. 30, p art II, 1876, pp. 1-12.
7 “ Spectrographic Analysis of Biological M aterial,” by J. Cholak
and R. V. Story, Industrial and Engineering Chemistry, analytical
edition, vol. 10, 1938, pp. 619-622.
8 “ Qualitative Spectrographic Analysis in the Arc, W ith Graphite
Electrodes,” by W. C. Pierce, O. R. Torres, and W. W. Marshall,
Industrial and Engineering Chemistry, analytical edition, vol. 12,
January 15, 1940, pp. 41-45.
9 Hilger Publication, no. 94/9, Adam Hilger, Ltd., London,
England, January, 1939.
10 “ Spectrographic M ethods of Trace Analysis,” by J. S. Owens,
Industrial and Engineering Chemistry, analytical edition, vol. 11, 1939,
pp. 59-63.

D eterm ination of A m m onia in
Condensed Steam
B y M. C. SCHWARTZ,1 W. B. GURNEY,* a n d T. E. CROSSAN3
Of the im purities which may affect the electrolytic
conductivity and pH o f water, the gases carbon dioxide and
am m onia are m ost resistant to removal. Carbon dioxide
has previously been given consideration alm ost to the exclusion
of the ammonia-removal problem. This second
paper in a series of three on the subject of purity of steam
by the authors deals with the preparation of am m oniafree
water by various m ethods, of which phosphoruspentoxide
batch distillation apparently gives the m ost
satisfactory results. The Nessler m ethod of am m onia
determination is also thoroughly discussed, as well as the
volumetric determ ination of am m onia with sodium hypobromite
and naphthyl red. The authors have used such
determinations o f am m onia in condensed steam to assist
in evaluating a suitable correction factor to com pensate
for its presence.
THE electrolytic-conductivity method of determining steam
purity is the one most frequently used in practice. Fitze
(l),1 Hecht and McKinney (2), Rummel (3), Powell (4, 5),
Place (6) have investigated this method. I t has the advantages
of extreme sensitivity and is susceptible to being made continuously
recording. It has the disadvantages of a need for a
correlating factor between electrolytic conductivity and soluble
solids, as well as correction factors for gases such as ammonia and
carbon dioxide which dissolve in the condensed steam to form
electrolytes. Rummel (7), Gurney (8), Powell and McChesney
(9) have developed apparatus designed to remove carbon dioxide
and ammonia from the sample without affecting the solids present
but, thus far, ammonia has resisted attempts at complete
removal, hence, a correction must be made for its presence. The
authors have investigated the determination of ammonia in
condensed steam to assist in evaluating the correction to be applied
for its presence.
P r e p a r a t io n o f A m m o n ia - F r e e W a t e r
The preparation of pure water has been investigated frequently
in the intervening years since 1894 when Kohlrausch and Heydweiller
(10) purified water to the extent that its electrolytic conductivity
was 0.04 micromho at 18 C and 0.054 micromho at 25 C.
The methods for the purification of water may be classified as
1 Research Chemist, Gulf States Utilities Company, and Research
Associate in W ater Technology, D epartm ent of Chemical Engineering,
Louisiana State University.
* Efficiency Engineer, Gulf States Utilities Company, B aton Rouge,
La. Mom. A.S.M.E.
5 Superintendent, Gulf States Utilities Company, B aton Rouge,
La. Mem. A.S.M.E.
4 Num bers in parentheses refer to the Bibliography at the end of
the paper.
Contributed jointly by the Gulf States Utilities Company, Louisiana
Station, Baton Rouge, La., and the Engineering Experim ent
Station, Louisiana State University, and accepted by the Joint Research
Committee on Boiler Feedwater Studies and presented at
the Spring Meeting, Boiler Feedwater Sessions, of T h e A m e r i c a n
S o c i e t y o f M e c h a n i c a l E n g i n e e r s , W orcester, Mass., M ay 1-3,
1940.
N o t e :
Statem ents and opinions advanced in papers are to be
understood as individual expressions of their authors, and not those
of the Society.
physical and chemical or a combination of both. Hibben (11)
classifies the physical methods of purification as falling into four
categories, namely, those employing vacuum, those employing
heat, those employing both vacuum and heat, and those employing
freezing. In his own experimental work (11), Hibben
uses the vacuum-sublimation process which includes all the processes
mentioned, since it employs heat, vacuum, and freezing.
The liquids to be freed from gas are vaporized and frozen in
vacuo. By chemical methods we refer to those cases in which
chemical substances, such as alkaline permanganate, bromine,
phosphoric acid, and similar reagents, are added to cause certain
impurities existing in the liquid to be retained in solution or to be
made volatile on the application of heat.
The most convenient methods for determining the purity of
water are the determination of the electrolytic conductivity and
that of the pH. The most resistant impurities to be removed from
water, which may affect the electrolytic conductivity and pH of
the water, are the gases, carbon dioxide, and ammonia, which, in
addition to physical solution in water, undergo chemical reaction
to form electrolytes as the bicarbonate and carbonate and ammonium
ion. Other gaseous impurities which may undergo similar
reaction, as sulphur dioxide, have not been considered in
reported work.
For some time carbon dioxide has been given consideration, to
the general exclusion of ammonia, as the gaseous impurity in
water to be removed. The probable reason for this has been due
to the indefiniteness of the electrolytic-conductivity method as
an indicator of purity, and to the lack of methods of determining
pH without adding electrolytes, which change the pH themselves,
as in colorimetric methods. The presence of carbon dioxide
in solution tends to lower the pH of pure water. When the glasselectrode
method for determining pH was applied to water considered
to be free from carbon dioxide, the pH was found to be
7.9 to 8 (12). The increase in pH above that of 7, the value for
pure water, was later ascribed to the presence of dissolved ammonia.
When definite measures were taken to insure the absence
of this impurity, the pH fell to 7 (13, 14).
Presumably the best procedure for preparing ammonia-free
water is by distillation, adding acidic substances, as P20 6 to retain
ammonium compounds in solution, and passing purified air
through the distillate (13, 14).
P r o c e d u r e f o r P r e p a r i n g A m m o n ia - F r e e W a t e r — B a t c h
D i s t i l l a t i o n W i t h P 20 6
Phosphorus pentoxide was added to 1 1 of laboratory-distilled
water. The solution was distilled at a rate of about 8 ml per min
in an all-Pyrex-glass still as shown in Fig. 1. The distillate was
collected in the conductivity-cell holder A from which it could be
collected for use or run to waste. Determinations of ammonia
in the distillate, as well as electrolytic-conductivity measurements,
indicated that about 200 ml of distillate should be discarded.
From then on a receiver could be placed at point B and the next
600 ml could be collected. Two such portions were collected and
1 1 of this water was then redistilled again, discarding the first
200 ml, and collecting for ammonia-free water the next 600 ml
of distillate. The flask from which the solutions were distilled
was equipped with a side tube and ground stopper which allowed
723

724 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
F i g . 1
D i s t i l l i n g A p p a r a t u s f o r P r e p a r a t i o n o f P u r e W a t e r
the vessel to be filled and emptied without dismantling the apparatus.
Passing purified air into the distillate was not done in
this work. A Leeds and Northrup No. 4866 mho-ohm conductivity
indicator with cell constant and temperature compensator
was used. The electrodes were platinized platinum.
Tables 1 to 8 and Fig. 2 indicate the comparative effectiveness
ofjVarious methods of preparing ammonia-free water and seem to
confirm the work of Ellis and Kiehl (13) and Cranston and Brown
(14) in establishing a phosphoric-acid distillation as the most
satisfactory method available.
It is interesting to note, as indicated by the data in Tables 1
to 8 and in other experiments where distillation has been carried
practically to completion, that as the distillation proceeds the
electrolytic conductivity decreases continuously. On the other
hand, in cases where chemicals have been added, the electrolytic
conductivity commences to increase when sufficient concentration
is reached to start carry-over.
The data of Table 8 indicate that simple distillation will not
drive off the ammonia completely, even when some 700 to 800
ml of an original 1 1 are distilled.
Considering the simplicity of the still used for preparing ammonia-free
water, in comparison with some of those described
in the literature, the electrolytic conductivity of the distillate

SCHWARTZ, GURNEY, CROSSAN—DETERM INATION OF THE PURITY OF STEAM 725
obtained is reasonably good. For comparison of the values obtained
with those of other investigators, the data of Table 9 are
included.
T A B LE 9 C O N D U C T IV IT Y OF
Prepared by
Thiessenn, H errm ann (15)
R im attei, P e tit (16)
Straub (17)
Best (18)
Bengough, S tu art, Lee (19)
Bencowitz, Hotchkiss (20)
K raus, D exter (21)
B ennett, Dickson (22)
Clevenger (23)
W eilana (24)
K endall (25)
Bourdillon (26)
K ohlrausch, Heydweiller (10)
(Pure w ater standing in contact with air)
‘P U R E ” W A T E R
Year M icromhos
1937 0 .0 6 5 -0 .0 8
1937 1.2
1935 0 .3 -0 .5
1929 0 .3 -0 .7 5
1927 0.045
1925 0 .0 6 -0 .0 7
1922 0 .0 5 -0 .1 2
1919
1919
0 .4
0.8
1918 0 .0 5 3-0.070
1916 0 .0 6 -0 .0 9
1913 0.2
1894 0 .0 5 -0 .1 1
0 .6 5 -0 .6 6
D e t e r m in a t io n o f A m m o n ia — N e s s l e r M e t h o d
Although Nessler proposed a reagent for the direct determination
of ammonia in 1856 (27), his method with various modifications
apparently remains the most widely used. Naturally, in
the intervening time a considerable volume of literature has accumulated,
but there are few other methods proposed for the determination
of small quantities of ammonia. Bjerrum (28) proposed
an acid titration for determining both ammonia and carbon
dioxide in distilled water. Makris (29) proposed a test using
tannin solution and silver nitrate which he claimed was more
sensitive than the Nessler test. Teorell (30) proposed a titrimetric
determination for minute amounts of ammonia by using the
reaction 2 NHS+ 3 NaBrO = 3 NaBr + 3 HjO + N,, the excess
NaBrO being determined iodometrically or by means of naphthyl
red, which is decolorized by NaBrO.
The most unfortunate feature of the Nessler method for determining
ammonia is the variation in the preparation of Nessler’s
reagent and its resulting varying sensitivity. In addition, the
reaction taking place is rather complex, and the Nessler reagent
undergoes change with time.
Nichols and Willits (31) indicate the reaction to be
2 (H gli.2 N al) + NH, = (NH,.H gI2) + 2 Nal)
2 (NH,.H gI2)___________ = NH2Hg2I, + N H J
2 (Hgl2.2 Nal) + 2 NH, = NH2Hg2I, + 4 N al + N H J
The compound NH2Hg2I, gives the color to Nesslerized ammonia
solutions.
Ammonia in the N essler R eagent
In their study of the Nessler reaction, Nichols and Willits (31),
working with a solution containing 2 ppm ammonia nitrogen, and
adding 5 ml Nessler reagent to a 100-ml sample, found that the
apparently clear yellow solution formed could be made clear and
colorless by passing it through a microfilter, using 0.06-gage cellophane
as the filtering medium, under a pressure of 290 lb of nitrogen.
With the exception of one other statement to this effect
quoted in Yoe (32), “J. H. Robertson and A. Hisey have recently
made an extensive experimental study on the nature of the
Nessler color and conclude that it is due to colloidally dispersed
particles in suspension and not in true solution (private communication);”
it is believed, in general, that an ammonia-free
Nessler reagent has a slight yellow-green color definitely distinguishable
from distilled water. In fact, the American Public
Health Association liquid standards for ammonia nitrogen call
for an addition of platinum salt (standard) to make the 0.00
standard. From the previously quoted statements, one would be
led to believe that the color in a Nessler reagent is due to the
presence of colloidal-like particles resulting from reaction with
the ammonia present in the water used to make the reagent.
In the limited time that was available for experimental work on
the filtration and ultrafiltration of a Nessler reagent, no success
resulted in trying to get a colorless solution, and the filtered
Nessler reagent gave results that were only 0.01 ppm lower than
those obtained with unfiltered Nessler reagent. Filtrations were
made through 12 sheets of quantitative filter paper and through
very thick mats of asbestos. Experiments with cellophane in a

726 TRANSACTIONS OF TH E A.S.M.E. NOVEMBER, 1940
Bilver-plated steel ultrafilter, using nitrogen, resulted in accelerating
the formation of a precipitate in the solution, instead of obtaining
further clarification. Normally a precipitate forms in a
Nessler reagent with the passage of time. Considering the complexities
arising from attempting to filter a strongly alkaline liquid
by ultrafiltration, it was decided not to attem pt further experimentation
along this line.
S e n s i t iv i t y o p t h e N e s s l e r R e a g e n t
Wirth and Robinson (33) point out that a zero blank does not
necessarily mean there is no ammonia present but, as they determined
experimentally using distilled water and the “standard
methods” reagent, that 0.02 ppm or less of ammonia nitrogen is
present. However, since they apparently prepared their ammonia-free
water by simple distillation, it is quite possible that
their minimum of 0.02 is actually somewhat higher.
E x p e r im e n t a l R e s u l t s
For the purpose of simplifying the determination of ammonia
by the Nessler method, La Motte liquid standards and reagents
were tried. Comparison of solutions was made in a La Motte-
Nessler tube comparator. Ammonia-free water, an ammoniumchloride
standard, and a Nessler reagent were prepared for comparison
of results.
Using 1 ml of Nessler reagent per 50-ml sample and a 10-min
reaction time as recommended by the American Public Health
Association Standard Methods of Analysis (34), the results seemed
definitely to limit the sensitivity of the determination. Using our
ammonia-free water, it took 0.05 ppm of added ammonia nitrogen
to give any indication of ammonia when compared with the La
Motte standards. After considerable testing we decided to use
2 ml Nessler reagent per 50-ml sample and compare after 30 min.
Tests were made in Nessler tubes, which were then closed with
rubber stoppers. We used ordinary Nessler tubes after we had
tried Nessler tubes as advocated by Powell (4) and found that
we obtained the same results in both types of tubes. Where the
sample to be collected was freely flowing, we caught the solution
in an open Nessler tube by allowing it to flow under constant
head into the bottom of the tube and then to overflow until the
sample was ready to be tested. A flow of about 90-ml per min,
for 30 min, was used. For accurate work, we checked and prepared
anew, if necessary, the ammonia-nitrogen liquid standards
used with a given Nessler reagent. It is quite simple to add
ammonia nitrogen to ammonia-free water to make a series of
standards and we prefer this procedure.
E l e c t r o l y t ic - C o n d u c t iv it y C o r r e c t io n f o r t h e P r e s e n c e
o f A m m o n ia i n C o n d e n s e d S t e a m
The “ammonia correction” was determined by actual fullplant
test. The data are presented in Table 10 and shown
graphically in Fig. 3. Ammonia was introduced into the
boiler-feedwater suction.
The slope of the line obtained in Fig. 3 gives 1 ppm ammonia
nitrogen = 8 micromhos. Extrapolation to zero ammonia gives
an extremely low electrolytic conductivity for gas-free condensed
steam. From the pH value of the condensate and the known
equilibrium existing between the bicarbonate and carbonate ions,
it is probable that the ammonia exists in solution in combination
with the bicarbonate and carbonate ion.
D e t e r m i n a t io n o f A m m o n ia V o l u m e t r ic a l l y W it h
S o d iu m H y p o b r o m it e a n d N a p h t h y l R e d
In order to check the results obtained by the Nessler method
for the concentration of ammonia in condensed steam, we have
considered other methods of analysis. As a result, we have investigated
the oxidation of ammonia by sodium hypobromite,
using naphthyl red as an indicator (30, 35). The method has
demonstrated excellent possibilities but, as yet, has not been
applied to routine testing. The reactions taking place are:
2 NH, + 3 NaBrO = 3 NaBr + 3 HaO + N2
NaBrO + naphthyl red = NaBr + colorless product
The reaction is very sensitive; 1 ml 0.0005 N naphthyl red being
equivalent to approximately 1.9 gamma ammonia nitrogen (1
gamma = 0.001 mg).
The procedure as developed by the authors for use in determining
ammonia in condensed steam is as follows:
Reagents (use ammonia-free water):
0.1 N NaOBr—2.5 g NaOH, 1.25 ml bromine, dilute to 500 ml
2 N Na2COj—106 g per 1
0.001 N NaOBr—dilute 0.1 N NaOBr using 5 ml 2 N Na2CO»
per 500 ml
4 per cent HBr—24 ml 48 per cent HBr, dilute to 200 ml
(NH

SCHWARTZ, GURNEY, CROSSAN—DETERM INATION OF TH E PURITY OF STEAM 727
concentration from naphthyl-red factor as obtained
under standardization; 1 gamma per 10 ml = 0.1 ppm.
Standardization:
A Take 10-ml sample of ammonia-free water and follow
procedure
B To 1 ml ammonia-nitrogen standard add 9 ml ammoniafree
water and follow procedure.
Sample calculation:
a 10 ml ammonia-free water require 2.10 ml naphthyl-red
solution
b 1 gamma ammonia nitrogen requires 1.60 ml naphthyl-red
solution
0.5 ml naphthyl-red solution is equivalent to 1 gamma ammonia
nitrogen
1 ml naphthyl-red solution = 2 gamma ammonia nitrogen.
The time required for a titration is only a m atter of a few
minutes. The standardization of the naphthyl-red-ammonianitrogen
factor should be made at least every day because the
solutions are not stable. Since the standardization is so easily
made, our practice has been to make it twice daily if tests are
made during an entire day. The constant check obtained by
frequent standardization serves a useful purpose. Our results
have shown that the method is good in the region 0.1 to 0.3
ppm ammonia nitrogen. In that range results are reproducible
to within 0.02 ppm or less, as shown in Table 11. For solutions
containing 0.5 gamma or 0.05 ppm ammonia nitrogen, the
naphthyl-red factor deviates from the constancy shown at higher
concentrations. Krogh (35) states that it is advisable to reduce
the volume of hypobromite used when working regularly with
amounts of ammonia nitrogen less than 1 gamma.
Ammonia
nitrogen
1 gam m a
2 gam m a
3 gam m a
T A B L E 11 T E S T R E SU L T S
N aphthyl-red-solution factor in gam m a
P pm »-------------- -am m onia nitrogen--------------- >
0 .1 1 .7 9 ,1 .8 8 , 2 .0 0 ,1 .8 5 1.88 avg
0 .2 1 .8 0 ,1 .9 6 ,1 .9 0 1.89 avg
0 .3 1 .9 6 ,1 .9 0 ,1 .9 1 1.92 avg
A c k n o w l e d g m e n t
The authors wish to thank H. C. Leonard, vice-president of the
Gulf States Utilities Company and Dean L. J. Lassalle, director
of the Engineering Experiment Station, Louisiana State University
for permission to publish this material. Grateful acknowledgment
is made to E. B. Powell, consulting engineer of
Stone and Webster Engineering Corporation, and to S. T. Powell,
consulting chemical engineer, for their advice, highly valued suggestions,
and encouragement throughout the course of the work
being reported. Members of the chemical staff of the Gulf
States Utilities Company assisting in this research were L. Young
and J. C. Hill.
B IBLIOGRA PHY
1 “ Determ ination of M oisture in Steam by Electrical Conductivity,”
by M. E. Fitze, Power, vol. 68, no. 12, 1928, pp. 484-485.
2 “ Electrical Conductance M easurements of W ater and Steam,
and Applications in Steam Plants,” by M. H echt and D. S. Mc­
Kinney, Fuels and Steam Power, Trans. A.S.M .E., vol. 53, 1931,
FSP-53-11, pp. 139-159.
3 “ Estim ation of Solids in Steam by Conductivity,” by J. K.
Rummel, Industrial and Engineering Chemistry, analytical edition,
vol. 3, 1931, pp. 317-320.
4 “ Steam Contam ination,” by S. T. Powell, Combustion, vol. 9,
no. 3, September, 1937, pp. 36-40; no. 4, October, 1937, pp. 27-31;
no. 5, November, 1937, pp. 25-31.
5 “ Design and Development of Apparatus for M easurement of
Steam Quality by Electrical C onductivity M ethods,” by S. T.
Powell, H . E. Bacon, Jr., I. G. McChesney, and F. Henry, Trans.
American Institute of Chemical Engineers, vol. 33, 1937, pp. 116-138.
6 “ Testing Steam Condensate for Its Quality and P u rity ," by
P. B. Place, Combustion, vol. 9, March, 1938, pp. 25-28.
7 “ Fluid Treating Device,” U. S. P aten t no. 2,046,583, J. K.
Rummel, July 7, 1936.
8 “ Determ ination of P urity of Steam by the Electrolytio-
C onductivity M ethod,” by W. B. Gurney, M. C. Schwartz, and T.
E. Crossan, published on page 728 of this issue.
9 “ Steam-Testing M ethod and A pparatus,” U. S. P atent no.
2,146,312, S. T. Powell and I. G. McChesney, February 7, 1939.
10 “ Uber reines W asser,” by F. Kohlrausch and A. Heydweiller,
Zeitschrift filr physikalische Chemie, vol. 14, 1894, pp. 317-330.
11 “ Removal of Dissolved Gases From Liquids by Vacuum Sublim
ation,” by J. H . Hibben, U. S. Bureau of Standards, Journal of
Research, vol. 3, 1929, pp. 97-104.
12 “ The Determ ination of the Hydrogen-Ion Concentration in
Pure W ater by a M ethod for Measuring the Electrom otive Force of
Concentration Cells of High Internal Resistance," by H. T. Beans
and E. T. Oakes, Journal of the American Chemical Society, vol. 42,
1920, pp. 2116-2131.
13 "T he Purification of W ater and Its pH Value,” by S. B.
Ellis and S. J. Kiehl, Journal of the American Chemical Society, vol.
57, 1935, pp. 2145-2149.
14 “ The pH of Distilled W ater, and the M easurement of the
Hydrolysis of Ammonium Sulphate, N itrate, Chloride, and A cetate,”
by J. A. C ranston and H . F. Brown, Trans. Faraday Society, vol. 33,
1937, pp. 1455-1458.
15 “ Eine Einfache M ethode zur Herstallung von Leit fahigkeits
wasser hochsten Reinheitsgrades,” by P. A. Thiessen and K. Herrmann,
Zeitschrift filr Elektrochemie, vol. 43, 1937, pp. 66-69.
16 “ Laboratory Preparation of Pure W ater,” by F. R im attei and
J. Petit, Bulletin de la Socii'tc-Chimie Biologique, vol. 19, 1937, pp.
1129-1133.
17 “ Continuous Production of Distilled W ater Free From Carbon
Dioxide and Ammonia,” by F. G. Straub, Industrial and Engineering
Chemistry, analytical edition, vol. 7, 1935, pp. 433-434.
18 "Acid Reaction and Carbon-Dioxide C ontent of Conductivity
W ater,” by R . J. Best, Australian Journal of Experimental Biology
and Medical Science, vol. 6, 1929, pp. 107-110.
19 "T he Routine Preparation of Low-Conductivity W ater,” by
G. D. Bengough, J. M. Stuart, and A. R. Lee, Journal of the Chemical
Society of London, part II, 1927, pp. 2156-2161.
20 "T h e Preparation of C onductivity W ater,” by I. Bencowit*
and H. T. Hotchkiss, Jr., Journal of Physical Chemistry, vol. 29, 1925,
p p .705-712.
21 “ An Im proved Still for Producing Pure W ater,” by C. A.
K raus and W. B. Dexter, Journal of the American Chemical Society,
vol. 44, p a rt 2, 1922, pp. 2468-2471.
22 “ The Bourdillon W ater Still,” by J. P. B ennett and J. G.
Dickson, Science, vol. 50, 1919, pp. 397-398.
23 “ R apid and Convenient M ethod for the Preparation of Conductivity
W ater,” by C. B. Clevenger, Industrial and Engineering
Chemistry, vol. 11, 1919, pp. 964-966.
24 “ The Equivalent Conductance of Electrolytes in Dilute Aqueous
Solution,” by H . J. W eiland, Journal of the American Chemical
Society, vol. 40, 1918, pp. 131-150.
25 “ Review. The Preparation of Conductivity W ater,” by J.
Kendall, Journal American Chemical Society, vol. 38, 1916, p. 2460.
26 “ The Preparation of Conductivity W ater,” by R. J. Bourdillon,
Journal of the Chemical Society of London, vol. 103, p a rt 1 ,1913,
pp. 791-795; Proc., vol. 29, pp. 124-125.
27 “ Uber das Verhalten des Jodquecksilbers und der Quecksilberverbindungen
iiberhaupt zu Ammoniak und uber eine neue R eaktion
auf Am m oniak,” by J. Nessler, thesis, Freiburg i. B., 1856 or “ On
the Behavior of Iodide of M ercury W ith Ammonia, and on a New Test
for Ammonia,” by J. Nessler, Chemical Gazette, 1856, vol. 14, p. 446.
28 ‘ ‘The Titrim etric Determ ination of Small Q uantities of Carbon
Dioxide and Ammonia in Distilled W ater,” by N. Bjerrum , Annales
Academiae Scientiarum Fennicae (A), vol. 29, 1927, 18 pp.
29 “ New Colorimetric Determ ination of Ammonia,” by K. G.
Makris, Zeitschrift filr analytische Chemie, vol. 84, 1931, pp. 241-242.
30 ‘ ‘A New M ethod for the Titrim etric Determ ination of M inute
Am ounts of Ammonia,” by T. Teorell, Biochemische Zeitschrift, vol.
248, 1932, p p . 246-255.
31 “ Reactions of Nessler’s Solution;” by M. L. Nichols and C. O.
Willits, Joum al of the American Chemical Society, vol. 56,1934, p. 769.
32 “ Photom etric Chemical Analysis,” by J. H . Yoe, vol. 1, J.
Wiley and Sons, Inc., New York, N. Y., 1928.
33 “ Photom etric Investigation of Nessler Reaction and W etting
M ethod for Determ ination of Ammonia in Sea W ater,” by H . E.
W irth and R . J. Robinson, Industrial and Engineering Chemistry,
analytical edition, vol. 5, 1933, pp. 293-296.
34 “ Standard M ethods for the Exam ination of W ater and Sewage,”
eighth edition, American Public H ealth Assn., New York, 1936.
35 “ M ethod for the Determ ination of Ammonia in W ater and
Air,” by A. Krogh, Biological Bulletin, vol. 67, 1934, pp. 126-131.

D eterm ination of P u rity of Steam by the
E lectrolytic-C onductivity M ethod
By W. B. GURNEY,1 M. C. SCHWARTZ,2 a n d T. E. CROSSAN8
In order to assure correct boiler operation and to
evaluate the effect of changes in water treatm ent or boiler
design on purity o f steam being delivered, it is essential
to be able to measure any contam ination that may occur.
This paper, which is the last of three papers by the authors
on the subject of steam purity, gives the procedure for
m aking a practically continuous determ ination o f steam
purity by the only satisfactory m ethod thus far available.
In this m ethod the electrolytic conductivity o f the condensed
steam is measured and then converted to a solids
basis. The distillation and the gravimetric m ethods are
also discussed by the authors.
THERE are two important and practical reasons for measuring
the purity of steam: (o) To determine the extent of
carry-over or steam contamination during normal operation
so that correctness of operation may be maintained at all times
and any irregularity detected at once; (b) to evaluate the effect
of any changes in water treatment or of any changes in the mechanical
design of the boiler on the purity of steam being delivered.
Therefore, any method of measuring steam purity must
be capable of practically continuous determination in order to be
of real value. Thus far, the only method that is able to meet
this requirement is one in which the electrolytic conductivity of
the condensed steam is measured and thereupon converted to a
solids basis by means of the proper calculations and corrections.
If either operation, a change in water treatment or a change in
the mechanical design of the boiler, is such as to result in priming
or a condition approaching that state, the increase in electrolytic
conductivity of the condensed steam is sufficiently great to be
readily recognized as such. I t is only when conditions are such
that the steam contamination is quite low, less than 1 ppm, and
remains at nearly that value after various changes, that the
electrolytic-conductivity method and other methods encounter
difficulties in evaluating the results of the changes. I t is understandable
why there will be difficulty in securing the necessary
accuracy when it is observed that condensed steam, containing
0.5 ppm solids, is 99.9995 per cent pure.
1 Determine the cell constant of the conductivity cell with a
solution of potassium chloride at constant temperature. All
electrolytic-conductivity values are referred to 25 C or 77 F by
use of the temperature compensator on the conductivity meter.
2 Remove carbon dioxide and some ammonia from the steam
sample by passing it through the degasifier. If the pH of the
condensed-steam sample rises from the acid to the alkaline side
on degasification, removal of carbon dioxide is indicated.
3 Observe the temperature of the condensed-steam sample.
Adjust the conductivity meter to correct for the measured cell constant
and temperature. Measure the electrolytic conductivity
of the condensed steam.
4 Determine the carbon dioxide and ammonia in samples of
condensed steam collected at the same time.
5 Compute the conductivity correction due to the presence of
ammonia and carbon dioxide, if any; subtract these from the
measured conductivity.
6 Subtract the value 0.05 micromho, that of pure water,
from the measured conductivity.
7 Compute the total solids in the steam by applying the
proper factor to the corrected electrolytic conductivity of the
condensed-steam sample.
For measuring the electrolytic conductivity of the condensedsteam
samples, Leeds and Northrup Meter No. 4866 is satisfactory
as a portable instrument. The instrument is equipped with
cell-constant and temperature-compensator control dials. For
continuous recording of conductivity measurements, we use a
Leeds and Northrup meter with a range of 0 to 2 micromhos in
steps of 0.02 micromho. The adjustment for cell constant and
temperature, however, is manually controlled. I t is possible to
compensate for temperature changes automatically by using resistance
thermometers, equipped with auxiliary slide-wire drums
which can be connected in the bridge circuit of the recording conductivity
meter and thus correct for any variation from 77 F.
A glass dip-type conductivity cell, having a cell constant not
greater than 0.1 is used. The electrodes are platinized platinum.
Potassium-chloride solutions are universally accepted as standard
reference solutions for electrolytic measurements. Jones
and Bradshaw (l)4 have prepared potassium chloride as follows:
P r o c e d u r e f o r D e t e r m i n i n g S t e a m P u r i t y b y E l e c t r o l y t ic
M e t h o d
“The purest material available by purchase was recrystallized
several times with centrifugal drainage. Qualitative chemical
The following procedure has been employed for determining
tests for magnesium and sulphate (the most likely impurities)
steam purity by the electrolytic method:
gave negative results. The salt was fused in a platinum crucible,
poured into a platinum dish, and transferred to a closed
1 Efficiency Engineer, Gulf State Utilities Company, B aton Rouge,
La. Mem. A.S.M .E.
bottle while still hot. The salt showed no tendency to gain in
* Research Chemist, Gulf States Utilities Company, and Research weight by absorption of moisture when allowed to stand on the
Associate in W ater Technology, D epartm ent of Chemical Engineering,
Louisiana State University.
balance.” Jones and Bradshaw (1) have considered the preparation
of potassium-chloride reference solutions and, as a result
* Superintendent, Gulf States Utilities Company, B aton Rouge,
La. Mem. A.S.M.E.
of painstaking work, have recommended the following value for
Contributed jointly by the Gulf States Utilities Company, Louisiana
Station, Baton Rouge, La., and the Engineering Experiment
an approximately 0.01 N solution, which is the most dilute solution
they considered. The solution is precisely defined as follows:
Station, Louisiana State University, and accepted by the Joint Research
Committee on Boiler Feedwater Studies and presented at the 0.745263 g of KC1 per 1000 g of solution in vacuum. At 25 C
Spring Meeting, Boiler Feedwater Sessions of T h e A m e r i c a n S o c i e t y the conductivity is 1408.77 micromhos. For preparing the standard
solution, Jones and Bradshaw recommend using conduc-
o f M e c h a n i c a l E n g i n e e r s , Worcester, Mass., M ay 1-3, 1940.
N o t e : Statem ents and opinions advanced in papers are to be
understood as individual expressions of their authors, and not those
of the Society.
4 Num bers in parentheses refer to the Bibliography at the end of
the paper.
728

GURNEY, SCHWARTZ, CROSSAN—DETERM INATION OF THE PURITY OF STEAM 729
tivity water prepared in contact with air. Such water will have
a specific conductance of approximately 1 micromho at 25 C.
The conductivity of the water used should be measured in advance
and subtracted from the measured conductivity of the solution.
I t is desirable to have available a more dilute standard
reference solution but, presumably, none has been prepared and
measured with the exactness with which Jones and his associates
have done for the more concentrated potassium-chloride solutions.
Powell (2) states that a 0.0001 N potassium-chloride
solution has a specific conductance of 15.27 micromhos at 25 C.
R e m o v in g C a r b o n D i o x id e a n d A m m o n ia F r o m C o n d e n s e d
S t e a m
The removal of carbon dioxide and ammonia from condensed
steam is difficult. Available methods have been considered
by the authors (3). An apparatus, represented in Fig. 1, has
been developed for this purpose. So far as the authors have been
Fio. 1 A r r a n g e m e n t o f A p p a r a t u s f o r R e m o v i n g C a r b o n D i­
o x i d e a n d A m m o n i a F r o m C o n d e n s e d S t e a m
able to determine by the use of this system, carbon dioxide is removed
to the extent that only undetectable amounts remain.
Removal of ammonia, however, is incomplete. The pH of the
condensed steam, measured with a glass electrode, is on the alkaline
side. Steam is sampled with the standard A.S.M.E. sampling
nozzle (4) installed in the main steam header close to the
boiler. From a steel shutoff valve it is then conveyed in stainless-steel
tubing to a tee where part of the steam passes directly
to a copper reboiling coil in the copper degasifying chamber.
From there the steam passes to a copper condenser. In addition
to the dissolved or suspended solids present, this condensedsteam
sample contains any soluble gases present in the steam and
is referred to as the straight-through sample. I t gives a prompt
indication of changes in steam composition. The sample passes
through a Pyrex-glass holder which contains a thermometer and
a conductivity cell.
The remaining portion of steam from the tee passes through a
stainless-steel condenser for partial cooling and from there into
the copper degasifying chamber. Upon coming in contact with
the steam-heated reboiling coil, the condensed steam again boils
violently. Dissolved gases are flashed out and steam serves also
as a scrubbing medium to assist in removing soluble gases. A
copper reflux or vent condenser condenses any steam that might
escape along with the dissolved gases and thus eliminates any
increase in concentration in the degasifying chamber. A vent
and overflow line assist in maintaining a constant water level in
the degasifying chamber. The cover on the degasifying chamber
fits loosely.
The degassed condensed steam issuing from the degasifying
chamber is further cooled by passing it through another stainlesssteel
condenser from which it passes into a Pyrex-glass holder
containing a thermometer and conductivity cell. This sample is
referred to as the degassed sample.
The steam flow should be about 280 ml per min. The flow
from the reboiling coil should be about 150 ml per min. The
flow of degassed condensed steam should be about 80 ml per min.
These flows are approximate for 720-F steam. Steam at other
temperatures requires variation in flow through the reboiling coil.
The copper reboiling coil is made of Vi-in. tubing, 1 0 ‘ / j turns,
and l'/a in. diam. The copper reflux coil is made the same as the
reboiling coil. With respect to the materials suitable for conveying
steam and condensate with minimum contamination, it has
been our experience that stainless steel is most suitable for highpressure
high-temperature steam, and copper is satisfactory for
condensate if a suitable flow is maintained.
In testing for the presence of carbon dioxide in the condensed
steam, it is customary to use N/44 or N/50 sodium hydroxide and
a sample of at least 250 ml. One or two drops are usually sufficient
to give a pink color with phenolphthalein. Under these
conditions, the amount of carbon dioxide present is considered
negligible. For determining ammonia in condensate, we have
used both the colorimetric Nessler method and the volumetric
sodium-hypobromite-naphthyl-red method, described in another
paper (3). It is important to note that the relative-conductivity
corrections to be applied for residual carbon dioxide and ammonia
in the condensed steam are approximately in the ratio of
1 to 10 per ppm of each present. I t is thus evident that, not
only does ammonia affect the results more appreciably, but the
residual ammonia must be determined with an accuracy of 0.01
ppm.
Our experience with the degasifier has indicated that no detectable
amounts of carbon dioxide remain in the condensed
steam, therefore no correction to the measured electrolytic conductivity
has been applied in this respect. The value of 1 ppm
ammonia nitrogen = 8 micromhos (25 C), which, as we have determined
experimentally, agrees well with currently used values
(2).
To convert electrolytic-conductivity measurements to ppm
dissolved solids, the procedure has been as follows: Small increments
of boiler water are added to degassed condensed steam.
The total solids in the boiler water are determined gravimetrically.
Providing the concentration increments are kept small,
the linearity of the relationship between ppm solids and electrolytic
conductivity is assured. A typical calibration is shown
in Table 1.
T A B L E 1 T Y P IC A L C A L IB R A T IO N O F E L E C T R O L Y T IC -C O N ­
D U C T IV IT Y M E A S U R E M E N T S
Boiler w ater added,
ml
0.00
0.10
0.20
0.40
Boiler w ater 1601 ppm .
Solids added,
ppm
0.00
0.32
0.64
1.28
C onductivity,
micromhos
0.82
1.49
2.09
3.38
Fig. 2 represents the plotted results. The slope of the line
gives the factor for converting micromhos to ppm; in this case
it is 1 micromho = 0.5 ppm dissolved solids (25 C).
A typical calculation for degassed condensed steam follows:

730 TRANSACTIONS OF THE A.S.M.E. NOVEMBER, 1940
T A B L E 3 VALUES F O R TO T A L SO LID S IN C O N D E N S E D STEA M
BY EV A PO R A TIO N A N D W E IG H IN G
Dissolved
From
solids.
B ibliography (6) Condensed ppm
T able 2, No. 6(a) Superheated steam 0.13
T able 2, No. 6(b) Superheated steam 0.10
T able 2, No. 7 S aturated steam 0.05
T able 2, No. 8 S atu rated steam 0.05
F io . 2 R e l a t io n B e t w e e n P pm D is s o l v e d S o l id s a n d E l e c ­
t r o l y t ic C o n d u c t iv it y
1.37 micromhos, measured conductivity; 0.14 ppm ammonia nitrogen;
no carbon dioxide; dissolved solids in condensed steam
= (1.37 — 1.12 — 0.05) X 0.5 = 0.10 ppm.
D is t i l l a t i o n M e t h o d
Place (5) has recommended distilling a given volume of condensed
steam in a glass still and measuring the electrolytic conductivity
of the solution remaining in the distilling flask. The
portion to be distilled is that sufficient to remove the volatile impurities.
Then, providing no dissolved solids have been added
from the glassware during the distillation, the remaining liquid
will contain only the dissolved solids originally present in the
sample. Place, however, considered carbon dioxide as the only
gaseous impurity in the solution. As the authors’ tests have
shown, it is not possible completely to remove ammonia by simple
distillation. Ammonia should therefore be determined in the
solution remaining in the flask and the proper correction applied
for its presence. A typical distillation on a sample of degassed
condensate is given in Table 2 (it is not necessary to take a degassed
sample for distillation).
In another paper (6), certain values for total solids in condensed
steam, obtained by evaporation and weighing, are given. Examination
of the residue disclosed appreciable amounts of iron
which originally were presumably not in solution but in suspension
and hence would not affect the conductance of the sample.
More direct evidence was found by adding pure water to the
residues in the platinum dish and obtaining the increase in conductivity.
On the basis that 1 micromho = 0.5 ppm dissolved
solids, the results given in Table 3 were obtained. These values
are in line with those derived from more direct electrolytic-conductivity
methods.
A c k n o w l e d g m e n t
The authors wish to thank H. C. Leonard, vice-president of the
Gulf States Utilities Company and Dean L. J. Lassalle, director
of the engineering experiment station, Louisiana State University
for permission to publish this material. Grateful acknowledgment
is made to E. B. Powell, consulting engineer of Stone
and Webster Engineering Corporation, and to S. T. Powell,
consulting chemical engineer, for their advice, highly valued suggestions,
and encouragement throughout the course of the work
being reported. Members of the chemical staff of the Gulf
States Utilities Company, assisting in this research, were L.
Young and J. C. Hill.
BIBLIO G RA PH Y
1 “The M easurement of the Conductance of Electrolytes—V.
A Redeterm ination of the Conductance of Standard Potassium-
Chloride Solutions in Absolute U nits,” by G. Jones and B. C. Bradshaw.
Journal of the American Chemical Society, vol. 55, 1933, pp.
1780-1800.
2 "Steam Contam ination,” by S. T. Powell, Combustion, vol. 9,
no. 5, November, 1937, pp. 25-31.
3 “ Determ ination of Ammonia in Condensed Steam ,” by M . C.
Schwartz, W. B. Gurney, and T. E. Crossan, published on page 5 of
this preprint.
4 “ Determ ination of Quality of Steam ,” A.S.M .E. Power Test
Codes, January, 1931, Instrum ents and Apparatus, P a rt 11.
5 "Testing Steam Condensate for Its Quality and Purity,” by
P. B. Place, Combustion, vol. 9, no. 9, M arch, 1938, pp. 25-28.
6 "D eterm ination of the Purity of Steam by Gravimetric and
Spectrographic M ethods,” by M. C. Schwartz, W. B. Gurney, and T.
E. Crossan, published on page 719 of this issue.
D iscussion1
D . S. M c K i n n e y 2 a n d M a x H e c h t . 8 In papers Nos. 1 and 3
the authors have made valuable contributions to the analysis of
water of high purity, and to the application of conductivity methods
to such water. For the sake of completeness, and for a
better understanding of the technique used, some details should
be amplified. For example, the drying temperature, which determines
the degree of hydration of the weighed residue, is not
given. The method of protecting the residue from hydration or
carbonation during weighing is not described.
Examination of Table 1, in first paper, indicates that the
residue was weighed to 0.1 mg. In view of the small quantity of
material, it would seem advisable to weigh the residue to 0.01 mg.
McKinney, in a discussion4 of Ulmer’s paper describes a method
for securing this sensitivity at relatively low cost with an analytical
balance.
D i s c u s s io n o f T h i r d P a p e r
Although numerous investigators have measured contamination
in steam, few have considered the problem of proper sam-
1 This is a combined discussion of the series of three papers on the
general subject “ Determ ination of P urity of Steam ,” by M. C.
Schwartz, W. B. Gurney, and T. E . Crossan. These papers will be
referred to in the discussion as first, second, and third, corresponding
to the order in which they are published.
2 A ssistant Professor, Chemical D epartm ent, Carnegie Institute
of Technology, Pittsburgh, Pa.
s Pittsburgh, Pa.
4 “ Determ ination by the E vaporation M ethod of Small Amounts
of Dissolved Solids in W ater Such as Condensed Steam From Boilers,”
by R. C. Ulmer, discussion by D . S. M cKinney, Proc. A.S.T.M.,
vol. 39, 1939, p. 1228.

GURNEY, SCHWARTZ, CROSSAN—DETERM INATION OF TH E PURITY OF STEAM 731
pling. It should be emphasized that the sample is secured from a
heterogeneous system, i.e., steam containing suspended solid or
liquid particles. In order to secure a representative sample from
such a system, it is essential that the velocity of the steam entering
the openings of the sampling nozzle be identical with that of
the main body of the steam approaching the openings. If the
sample is withdrawn too rapidly, too great a proportion of steam is
obtained, if too slowly the proportion of suspended liquid or solid
will be too high. It is pertinent to ask if the authors have met
this requirement.
The writers wish again to call to the attention of those interested
in conductivity measurements, that the relation of total
solids to specific conductance is likely to be much more variable
than the relation of dissolved salts to specific conductance. The
authors apparently do not distinguish between the terms total
solids and dissolved salts as indicated by their method of calibration.
In Table 2 of the third paper, the authors determine gravimetrically
the total solids in one sample of boiler water to obtain
the relation between solids in the steam and conductivity. It
should be noted that an unknown fraction of the boiler solids
are nonconducting (Fe20 3, Si02, etc). Furthermore, the fractions
of nonconducting material may vary considerably during the
normal run of a boiler. Yet in Table 3, the same factor is used
to calculate the soluble portion of the evaporated residue from
steam samples, which was found to be only about one third of the
weight of evaporated residue.
It is suggested that the dissolved salts in the boiler water
should have been made the basis of the calibration. This point
was discussed extensively in an earlier paper6 by the writers
and re-emphasized in the discussion by Hecht of Ulmer’s paper. 6
Various factors for calculating dissolved substances in water
from conductivity measurements have been reported in the literature.
Reference to Table 9 of the writers’ paper7 shows that,
with the exception of H and OH, all common ions in water have
specific conductances per ppm between 1.14 micromhos for NO3
and 2.99 micromhos for Ca++. For OH- the value is 11.3 and
for H + 347 micromhos per ppm; all values being referred to 25 C.
It is obvious that variations in H and OH concentrations will
materially affect the relation of dissolved salts to conductances.
It should therefore be emphasized that the factor reported for
specific plants applies only to those plants. In general the appropriate
factor will depend upon the water supply and its treatment.
E. B. P o w e l l .8 The authors have demonstrated the practical
feasibility of power-plant measurement of solids carryover
in proportions as low as 0 . 1 ppm, and have described procedures
by which a highly creditable approach to precision can be
maintained in such measurements continuously. The electrolytic-conductivity
method obviously does not give a direct measure
of total solids to be encountered in condensed steam. Extensive
work on corrections and calibrations, given special impetus
by that of Hecht and McKinney, 6 now regarded as a classic in the
power-plant field, has been carried on more or less continuously
over the last 10 years. Statistical data are not yet available to
cover correction factors for all water conditions. However, by
securing appropriate calibrations for different water conditions to
be met, following the general procedures described by the present
authors, it is entirely practicable to set up data from which the
6 “Electrical-Conductance M easurements of W ater and Steam and
Applications in Steam Plants,” by D. S. M cKinney and Max Hecht,
Trans. A.S.M .E., vol. 53,1931, FSP-53-11, pp. 139-151.
‘ Discussion reference 4, p. 1229.
7 Discussion reference 5, p. 147.
8 Consulting Engineer, Stone & W ebster Engineering Corporation,
Boston, Mass. Mem. A.S.M.E.
individual plant can interpret the conductivity values in terms
of actual solids carry-over to the degree of precision mentioned.
The ability to measure is indeed an important factor in the ability
to control, and the solids in steam are no exception. The authors’
contribution will be decidedly welcome to the industry.
S. T. P o w e l l 9 a n d H. E. B a c o n .10 The steam-quality data
presented in this series of papers indicate the unusually low soluble-solids
concentration of about 0 . 1 ppm, yet they are confirmed
by conductivity, gravimetric, and spectrographic measurements.
The authors have certainly shown to the satisfaction of
the most critical reviewers that the total solids in solution in the
samples tested were less than 0.2 ppm. This demonstrates that,
if currently accepted specifications for steam quality are changed
in the direction of higher purity, it is entirely feasible to apply
tests which will give meaning to the specifications.
There can be no doubt that the more critical operation of highpressure
boilers requires a specific knowledge of steam quality,
which, to be of value, should cover the full range of possible operating
conditions. Each of the variables, such as boiler-water
concentrations, water levels, and steam flow should be changed
one at a time to develop a basis for predicting performance under
all conditions. This requires evaluation of the carry-over at frequent
intervals, or preferably continuously, which precludes the
use of gravimetric determination and leaves no other recourse
except estimating solids by conductivity methods. Qualitative
measurement, where carry-over is appreciable, presents no problem.
The major difficulties, and magnification of all inaccuracies,
become increasingly great with improved quality of steam.
Methods for measuring the purity of steam have primary significance
in performance tests covering a limited period, but their
utility for routine control of operations may be quite valuable.
Permanent and continuously recording steam-conductivity apparatus
has enabled the operators of one large plant to increase
boiler-water concentrations from a tentative limit of 2500 ppm
to 3500 ppm as the standard for controlling the plant. The
availability of instantaneous indications of carry-over may thus
permit the exploration of numerous changes in operation which
promise greater economy and efficiency. In such cases, the refinements
described by the authors and corrections for dissolved
gases are not practicable except for brief periods when highly accurate
tests are desired. However, the operators soon learn to
discount even large errors from such recognized sources and to
correlate the meter charts with actual operating conditions. At
the plant mentioned, the conductivity of the boiler blowdown is
also recorded to provide a check on boiler-water concentrations.
The experience of the writers has been almost entirely with the
vacuum type of steam-sample degasifier. The residual ammonia
left by this equipment is of the same order of magnitude as that
cited by the authors for the reboiling type of degasifier. The
vacuum apparatus, for which the practical application has been
described previously, 1 1 operates for long periods with very little
attention. I t would be of interest to have the authors suggest
the important details in using their apparatus for continuous
operation in routine plant control.
In testing condensed steam for ammonia in the atmosphere of
most steam-generating stations, it is almost impossible to make
and preserve large quantities of ammonia-free water. We have
found the use of Folin’s ammonia permutit to be satisfactory for
preparing water to be used in making up ammonium-chloride
8 Baltimore, M d. M em. A.S.M .E.
10 Special Representative, Stoker D epartm ent, Westinghouse
Electric & M anufacturing Company, Philadelphia, Pa. Mem.
A.S.M .E.
11 “Steam Contam ination—III. Determ ination of Steam Quality
,” by S. T. Powell, Combustion, vol. 9, no. 5, Nov., 1937, pp. 25-31.

732 TRANSACTIONS OF TH E A.S.M.E. NOVEMBER, 1940
standards. These are only used, however, in the adjustment of
permanent standards for each new batch of Nessler reagents.
We hold no brief for conductivity measurements as providing
absolute accuracy, realizing the deficiencies that exist. From
our experience, however, we are convinced that steam-station
operators now have available a means for closely estimating the
performance of the boiler with respect to steam quality. Further
improvement in apparatus for such measurements presents a
worth-while problem for future study.
A l f r e d W a t s o n . 12 In the application of the electrical-conductivity
method to the determination of the purity of steam, the
most perplexing problem has been the separation of the effect of
dissolved solids from that of the ionized dissolved gases, such as
ammonia and carbon dioxide. Various investigators have alternated
between the two main lines of attack: (1) Measurement of
the concentrations of ammonia and carbon dioxide, calculation of
the increment of conductivity due to these concentrations, subtraction
of this value from the observed conductivity, and calculation
of the residual conductivity into terms of dissolved salts;
and (2) physical elimination, in so far as possible, of dissolved
gases before measurement of the conductivity, followed by direct
calculation of the latter into terms of dissolved salts.
Application of the first method has been hampered by uncertainty
as to the corrections to be applied when both ammonia
and carbon dioxide were present in the sample of condensate;
application of the second by the difficulty of removing all the ammonia
by any practical means. The authors have offered the
compromise of removing substantially all of the carbon dioxide
and part of the ammonia, so that correction need be made only
for low residual concentrations of ammonia.
While equipment for degasification of condensed-steam samples
may well form part of the permanent testing equipment in a large
central station or industrial steam plant, it is inconvenient to
transport from plant to plant. The engineer engaged in field
tests is predisposed, therefore, to rely upon the first method of
leaving the ammonia and carbon dioxide in the sample, and to
develop as accurate a means as he can for correcting his observed
values for conductivity in the light of his analyses for these gases.
Offhand, it might seem suitable in making correction for carbon
dioxide and ammonia to read off from the specific-conductivity
curves of each, respectively, that conductivity value corresponding
to the determined concentration, to add these two values, and
to subtract the total from the observed conductivity of the sample
regarding the difference as the conductivity due to dissolved
solids.
In a great many instances quite anomalous values have been
obtained through the use of this method of correction. The most
common anomaly is that of a negative value when the total correction
is subtracted from the observed conductivity.
Such a value is, of course, an impossibility.
Wherein does the error lie
It is the writer’s opinion that the method often
followed of making separate corrections for ammonia
and for carbon dioxide and then totaling
these two values is incorrect, because it neglects
the fact that the sample of water can have one
and only one pH value and one equilibrium between
the various ions for the conditions which
existed when the conductivity was measured.
Actually, since ammonia is a base and carbon
dioxide an acid in solution, with different relative
amounts of these two, the relative concentrations
of H+, NH,+, O H -, HCOj- , and CO,—
ions in the sample will vary. The specific conductivity of OH~
ion is, however, nearly 3 times, and that of H+ ion 5 times that
of NH

GURNEY, SCHWARTZ, CROSSAN—DETERMINATION OF THE PURITY OF STEAM 733
The curves of Fig. 1 are expressed in terms of ammonia and of
“free” rather than total carbon dioxide, as determined by standard
methods. The most significant feature is the extremely
slight effect of carbon dioxide upon the conductivity when ammonia
is also present and the pH value is below 7.
The suggested procedure for testing a sample of steam condensate
is as follows:
1 (a) The concentrations of “free” carbon dioxide and ammonia
in the sample are determined by recognized standard
methods.”
(6) The pH value of the sample may be determined by the
glass electrode, as with Coleman electrometer Model 3D, or Leeds
and Northrup No. 7661-A1, or equal.
2 In case the content of “free” carbon dioxide is zero or negative,
or the pH value exceeds 7, for the most accurate work it will
be advisable to adjust the “free” carbon dioxide to an adequate
excess or the pH to below 7 by means of pure carbon dioxide, and
recheck the ammonia content.
3 The specific conductivity of the sample under these conditions
is obtained by means of a conductivity-resistivity indicator,
such as Leeds and Northrup No. 4866, or equal.
4 The correction for dissolved gases is read from a chart of the
type of Fig. 1 of this discussion, and subtracted from the observed
conductivity.
5 The corrected conductivity is converted into ppm of dissolved
salts by the use of a suitable factor.
Table 1 with this discussion, presents data on conductivity,
ammonia, and carbon dioxide, obtained in a plant test on steam
quality. In this case columns 5 and 6 give the correction values
and the dissolved solids, respectively, when separate correction
factors for carbon dioxide and ammonia were used, neglecting the
fact that the sample could only have one pH value. Columns 7
and 8 give the values obtained when the correction value was read
off from the curves of Fig. 1. It is notable that in the former case
the value for dissolved solids is negative in every case but one—
an obvious impossibility. In the latter all values are on the positive
side.
The suggestion of adding carbon dioxide to a sample of steam
condensate instead of removing it may seem unorthodox, but is
as fundamentally sound as the data for the dissociation of carbonic
acid and ammonium hydroxide. It should be kept in mind
that the curves of Fig. 1 are calculated from these values in the
ls “Standard M ethods for the Exam ination of W ater and Sewage,”
Eighth edition, American Public H ealth Association, New York,
N. Y., 1936.
literature, and the job of verifying them completely in practical
determinations has only begun. We plan much further work
along this line, and, finally, to present this entire subject in more
complete form at a later date.
A u t h o r s ’ C l o s u r e
As Messrs. E. B. and S. T. Powell point out in their discussions,
it has been demonstrated to be possible to determine, continuously,
solids carry-over in steam in proportions as low as 0.1
ppm. The authors feel that they have contributed to this important
power-plant problem. The variety of conditions found
in different plants and the requirements of different individuals,
as pointed out by Mr. Watson, seems to indicate that considerable
flexibility will be required in test methods. Under such circumstances,
the problem of carry-over measurement must be
attacked from all angles.
The drying temperature for gravimetric determination of total
solids residue, requested by Messrs. McKinney and Hecht was
110 C. The weight of the residue upon successive weighings,
involving the incidental transferring from desiccator to balance,
did not change, indicating that hydration and carbonation were
not taking place. I t is possible to weigh the residue to 0.01 mg
but the authors do not believe this accuracy is warranted at present.
It is worth-while to point out the fact that the temperature
between successive weighings should be practically constant.
The question of steam sampling has been considered by the
authors and suitable flow maintained to secure proper velocities.
This phase of the problem has also been emphasized by S. T.
Powell.11 Messrs. McKinney and Hecht correctly indicate that
dissolved solids rather than total solids should be made the basis
of calibration for electrolytic conductivity versus dissolved solids.
In the case of the Louisiana Station, the dissolved and total solids
are substantially the same. The correspondence between
gravimetrically determined dissolved solids, after subtracting
the weight of nonconducting residue, and independently determined
dissolved solids by conductivity measurement bears this
out.
Mr. Watson’s procedure for computing the correction to be
applied to electrolytic-conductivity measurements by determining
the ammonia and carbon-dioxide concentration in the condensed
steam, adjusted to a pH equal or less than 7 with carbon
dioxide, is quite interesting.
In conclusion the authors would again like to repeat: The
problem of producing pure steam is of great importance and justifies
the attention of all power-plant operators.