Alfred Still.

Principles of electrical design; d. c. and a. c. generators online

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peratures in order to expel the superfluous varnish. Mica is
hard and affords good protection against mechanical injury;
but it is not suitable for insulating corners or surfaces of irregular

3. Treated Fabrics, such as Varnished Cambric; Empire
Cloth, etc. These provide a means of applying a good insulat-
ing protection to coils of irregular shape. Linseed oil is very
commonly used in the preparation of these insulating cloths
and tapes because it has good insulating properties and remains
flexible for a very long time.

It is common practice to impregnate the finished armature
coil with an insulating compound, and press it into shape at a
fairly high temperature. When this is done, ordinary untreated
cotton tape is used in place of the varnished insulation.

Fig. 32 shows a typical slot lining for a 500-volt winding.
This can, however, be modified in many respects, and^the reader



Hard Wood Wedge

(Slot Lining of Press-
N pahn about 0.010 "
< Thick

v Cotton Tape; Empire
Xiloth or Micanite to
Total Radial Thick-
ness about 0.030"
Cotton Covering;
Braiding or Tape on
Conductors, of 0.014"
to 0.025"Single

\Pres8pnhn or Fish
Paper Dividing Strip,
of about O.Olo'Tbick

h Paper at Bottom
of Slot about 0.020 >J

need not at present concern himself with practical details of
manufacture. It is obvious that the insulation should be so
arranged as to leave the greatest possible amount of space for

the copper. The insulation
may be placed around the in-
dividual coils, or in the slot
before the coils are inserted.
If preferred, part of the insu-
lation may be put around the
coils and the remainder in
the form of a slot lining. The
essential thing is to have
sufficient thickness of insula-
tion between the cotton-cov-
ered wires and the sides of the
slot. The following figures
FIG. 32. Insulation of conductors in may be used in determining

the necessary thickness of slot

lining. These figures give the thickness, in inches, of one side
only, and this is also the thickness that should be provided be-
tween the upper and lower coil sides in the slot.

For machines up to 250 volts . 035 in.

For machines up to 500 volts . 045 in.

For machines up to 1,000 volts 0. 06 in.

For machines up to 1,500 volts 0.075 in.

In high-voltage machines an air space is sometimes allowed
between the end connections, i.e., the portions of the coils not
included in the slots. This air clearance would be from % to
K in. for a difference of potential of 1,000 volts, with an addition
of y$ in. for every 500 volts. In regard to the surface leakage
where the coils pass out from the slots; a breakdown of insulation
at this point is usually guarded against by allowing the slot
lining to project at least J- in. beyond the end of the slot. For
working pressures above 500 volts, add J^ in. for every additional
500 volts.

The finished armature should withstand certain test pressures
to ensure that the insulation is adequate. The standardization
rules of the A.I.E.E. call for a test pressure of twice the normal
voltage plus 1,000 volts.

29. Current Density in Armature Conductors. The per-
missible current density in the armature coils is limited by tern-


perature rise. The hottest accessible part of the armature,
after a full-load run of sufficient duration to attain very nearly
the maximum temperature, should not be more than 40 or 45C.
above the room temperature. No definite rules can be laid
down in the matter of armature conductor section because the
ventilation will be better in some designs than in others, and a
large amount of the heat to be dissipated from the armature core
is caused by the iron loss which, in turn, depends upon the flux
density in teeth and core.

The current density in the armature windings generally lies
between the limits of 1,500 and 3,000 amp. per square inch. If
the armature were at rest, the permissible current density would
be approximately inversely proportional to the specific loading,
q, i.e., to the ampere-conductors per inch of armature periphery.
When the armature is rotating, the additional cooling effect due
to the movement through the air will be some function of the
peripheral velocity, and, for speeds up to about a mile a minute
or, say, 6,000 ft. per minute the permissible increase of current
density will be approximately proportional to the increase in
speed. The constants for use in an empirical formula expressing
these relations are determined from tests on actual machines,
and the writer proposes the following formula for use in deciding
upon a suitable current density in the armature winding:

where A stands for amperes per square inch of copper cross-
section, and v is the peripheral velocity in feet per minute.

30. Length and Resistance of Armature Winding. Before the
resistance drop and the PR losses in the armature can be calcu-
lated, it is necessary to estimate the length of wire in a coil.
This length may be considered as made up of two parts: (1) the
"active" part, being the straight portion in the slots, and (2)
the end connections.

The appearance of the end connection is generally as shown in
Fig. 33, and since the pitch of the coil is measured on the cir-
cumference of the armature core, the sketch actually represents
the coils laid out flat, before springing into the slots.

The angle a which the straight portion of the end connections

makes with the edge of the armature core is sin" 1 > where
X is the slot pitch; s, the slot width; and 5, any necessary clear-




ance between the coils. This clearance need be provided only
in high-pressure machines, or when it is desired to improve
ventilation. For approximate calculations on machines up to
600 volts, the angle a may be calculated from the relation

sin a

In practice the angle a usually lies between 35 and 40 degrees.
The length of the straight part AC (Fig. 33) is where BA


T .

is half the coil pitch, or x in the case of a full-pitch winding.

FIG. 33. End connections of armature coil.

The portion of the end connections between the end of the
slot and the beginning of the straight portion AC is about %
in. in low-voltage machines, increasing to 1 in. in machines
for pressures between 500 and 1,000 volts. The allowance for
the loop where the coil is bent over to provide for the lower half
of the coil clearing the upper layer of conductors will depend
upon the depth of the coil-side and therefore upon the depth of
the slot. If d is the total slot depth, in inches, an allowance equal
to 2d will be sufficient for this loop. The total length of coil
outside the slots of a low-voltage armature will therefore be

where r must be taken to represent the coil pitch instead of the
pole pitch if the winding is of the chorded or short-pitch type.


The average length per turn of one coil will be 2l a -f l e where l a
is the gross length of the armature core. A small addition should
be made for the connections to the commutator, especially if
the coil has few turns. 1 The resistance of each coil, and there-
fore of each electrical path through the armature, may now be
readily calculated. In arriving at the resistance of the armature
as a whole it is important to note carefully the number of coils
in series in each armature path, and the number of paths in
parallel between the terminals of the machine. As a check on
the calculated figures, the IR drop, or the PR loss, in the arma-
tures of commercial machines, expressed as a percentage of the
terminal voltage or of the rated output, as the case may be, is
usually as stated below:

In 10-kw. dynamo 3 . 1 to 3 . 8 per cent.

In 30-kw. dynamo 2.6to3.2 per cent.

In 50-kw. dynamo 2.4to3.0 per cent.

In 100-kw. dynamo 2 . 1 to 2 . 6 per cent.

In 200-kw. dynamo 2.0 to 2.4 per cent.

In 500-kw. dynamo 1 . 9 to 2 . 1 per cent.

1 There is another type of end connection, known as the involute end
winding. It is not much used ; but those interested in the matter are referred
to the first volume of "The Dynamo " by HAWKINS and WALLIS ( WHITTAKER
& Co.), where the manner of calculating the length of these end con-
nections is explained.



31. Hysteresis and Eddy-current Losses in Armature Stamp-
ings. The loss due to hysteresis in iron subjected to periodical
reversals of flux may be expressed by the formula,

watts per pound = K h B l -*f

where KH is the hysteresis constant which depends upon the
magnetic qualities of the iron. The symbols B and / stand as
before for the flux density and the frequency.

An approximate expression for the loss due to eddy currents
in laminated iron is,

watts per pound = K e (Bft)*

where t is the thickness of the laminations, and K e is a constant
which is proportional to the electrical conductivity of the iron.

With the aid of such formulas, the hysteresis and eddy-current
losses can be calculated separately and then added together to
give the total watts lost per pound. This method will give good
results in the case of transformers; but when the reversals of
flux are due to a rotating magnetic field, as in dynamo-electric
machinery, the losses do not follow the same laws as when the
flux is simply alternating; and moreover there are many causes
leading to losses in built-up armature cores which cannot easily
be calculated. These additional losses include eddy currents due
to burrs on the edges of stampings causing metallic contact
between adjacent plates. There are also eddy currents produced
in the armature stampings due to the fact that the flux cannot
everywhere be confined to a direction parallel to the plane of
the laminations. Some flux enters the armature at the two
ends and also into the sides of the teeth through the spaces pro-
vided for ventilation. Since this flux enters the iron in a direction
normal to the plane of the laminations it is sometimes account-
able for quite appreciable losses. For these reasons calculations



of core losses should be based on the results of tests conducted
with built-up armatures rotated in fields of known strength.
Such tests are made at different frequencies, and the results,
plotted in graphical form, give the total watts lost per pound of
iron stampings at different flux densities, a separate curve being
drawn for each frequency. The reader is referred to the hand-
books of electrical engineers for useful data of this sort; but for
approximate calculations of core losses, the total iron loss per
cycle may be considered constant at all frequencies. This
assumption allows of a single curve being plotted to show the
connection between watts lost per pound and the product kilo-
gausses X cycles per second. This has been done in Fig. 34
which is based on experiments conducted by MESSRS. PARSHALL
and HobART and confirmed lately by PROFESSORS ESTERLINE
and MOORE at Purdue University. The curve gives average
losses in commercial armature iron stampings 0.014 in. thick.
Great improvements in the magnetic qualities of dynamo
and transformer iron have been brought about during the last
20 years, and the introduction of 3 to 4 per cent, of silicon in
the manufacture of the material known as silicon steel has
given us a material in which not only the hysteresis, but also
the eddy-current losses, have been very considerably lowered.
There are great variations of quality in armature stampings,
and values obtained from Fig. 34 would not be sufficiently
reliable for the use of the commercial designer of any but small
machines. By taking pains in assembling in stampings to
avoid burrs and short-circuits between adjoining plates, the
total iron loss may be considerably reduced. In large machines,
with a surface which is small in proportion to the volume,
the losses will usually be less than would be indicated by Fig.
34. In the absence of reliable tests on machines built with a
particular quality of iron punchings, it is suggested that the
values obtained from Fig. 34 may be reduced as much as 50 per
cent, in cases where extra care and expense with a view to re-
ducing losses are justified; and for silicon steel (a more costly
material than the ordinary iron plates) the reduction may be as
much as 70 per cent.

When calculating the watts lost in the armature core, it is
necessary to consider the teeth independently of the section
below the teeth. This is because the flux density in the teeth
is not the same as that in the body of the armature.


The calculation of the watts lost in the core below the teeth
is a simple matter provided the assumption can be made that
the flux density has the same value at all points. Although in-
correct, this assumption is very commonly made; and, for the
purpose of estimating the rise in temperature, the flux density
may be calculated by dividing half the total flux per pole by the
net cross-section of the armature core below the teeth. A
reference to Fig. 34 will give the watts per pound, and this,




1 9



^ 6













' '



For Carefully Assembled
High Grade Iron Stampings
Multiply the Watts Obtained
from Curve by 0.5

For Silicon Steel Stampings
Multiply the Watts Obtained
from Curve by 0.3











100 200 300 400 500 600 700 800 900 1000 1100
Kilogausses x Frequency, or ^-

FIG. 34. Losses in armature stampings.

when multiplied by the total weight of iron in the core (excluding
the weight of the teeth), will be the approximate total loss due
to hysteresis and eddy currents.

In order to calculate the losses in the armature teeth, it is
necessary to know exactly what is the flux density at all sections
of the tooth. This is not readily calculated, because some of the
flux from the pole pieces enters the armature through the sides
of the teeth and the bottom of the slots. Again, in the case
of armatures of small diameter having teeth of which the
taper may be considerable, the change of section alters the flux
density and the degree of saturation, so that it is almost im-
possible to determine accurately the average value of the tooth


density for use in calculating the watts lost. The question of
flux density in the teeth will be again referred to when discussing
the m.m.f. necessary to provide the required flux; and formulas
will be developed for use in calculating the actual flux density
in the iron of the teeth. For the purpose of estimating the tem-
perature rise, we shall assume that the whole of the flux from
each pole enters the armature through the teeth under the pole
(the effect of fringing at pole tips being neglected); and if the
teeth are not of uniform section throughout their length, the
average section will be used for calculating the flux density.
Thus, let

3> = the total flux per pole,
r == pole pitch,
X = tooth pitch,
t = width of tooth at center,
l n = net length of iron in armature,

pole arc

r = ratio -^ rr-r-

pole pitch

then the number of teeth under each pole is r - and the flux per


tooth is The flux density in the tooth, on the assumptions

previously made, would, therefore, be -,- gausses if the dimen-

TTlL n

sions are expressed in centimeters. By referring to the curve
Fig. 34, the watts lost in the teeth per pound of iron can be found.
The length l n is simply the gross length of the armature core less
the space taken up by ventilating ducts and insulation between
armature stampings. The question of vent ducts will be taken
up immediately; but, even when a suitable allowance has been
made for the spaces between the assembled sections of the
armature core, a further correction must be made to allow
for the thin paper or other insulation between the stampings.
The space taken up by this insulation will vary between 7 and
10 per cent, of the total space. Thus, if l a is the gross length
of. the armature, and l v the total width of all vent ducts, the
net length of the armature core would be.

l n = 0.92 (l a - l v )

if the space occupied by the insulation between laminations is
8 per cent.


32. Usual Densities and Losses in Armature Cores. The

flux density in the core below the teeth will be determined by
considerations of heating and efficiency. The same may be
said of the tooth density, but in this case the total weight of iron
is relatively small, and higher densities are permissible. It is
desirable to have a high flux density in the teeth because this
leads to a " stiff er" field and reduces the distortion of air-gap
flux distribution caused by the armature current. Better pres-
sure regulation is thus obtained, and also improved commutation,
especially on machines without interpoles where the fringe of
flux from the leading pole tip is used for reversing the e.m.f.
in the short-circuited coils under the brush. If the density in
the teeth is forced to very high values, the losses will be excessive,
especially if the frequency is also high; another disadvantage
being the large magnetizing force necessary to overcome the
reluctance of the teeth and slots.

The accompanying table gives flux densities in teeth and core
that are rarely exceeded in ordinary designs of continuous-current


Frequency, /

Density in teeth

Density in core













As a guide to the permissible losses in the armature punch-
ings of D.C. machines, the following figures will be useful.
They are based on modern practice and should not be greatly
exceeded if the efficiency and temperature rise are to be kept
within reasonable limits.

Output of machine Core loss, expressed as percentage

of output

10 kw 2. 8 to 3. 3

20 kw 2.5 to 3.0

50 kw 2.0 to 2. 4

100 kw 1.5 to 1.8

500 kw 1 . 3 to 1 . 5

1,000 kw 1 . 2 to 1 . 4


33. Ventilation of Armatures. Recent improvements in
dynamo-electric machinery have been mainly along the lines
of providing adequate means by which the heat due to power
losses in the machine can be carried away at a rapid rate,
thus increasing the maximum output from a given size of

Still air is a very poor conductor of heat; but when a large
volume of cool air is passed over a heated surface, it will effectu-
ally reduce the temperature which may, by this means, be kept
within safe limits.

The rotation of the armature of an electric generator will pro-
duce a draught of air which may be sufficient to carry away the
heat due to PR and hysteresis losses without the aid of a blower
or fan. Self-ventilating machines are less common at the present
time than they were a few years ago; but, by providing a suffi-
cient number of suitably proportioned air ducts in the body
of the armature, machines of moderate size may still be built
economically without forced ventilation. Radial ducts are
provided by inserting special ventilating plates at intervals
of 2 to 4 in., and so dividing the armature core into sections
around which the air can circulate. The width of these venti-
lating spaces (measured in a direction parallel to the axis of
rotation) is rarely less than % in. or more than % in. in machines
without forced ventilation. A narrower opening is liable to
become choked up with dust or dirt, while the gain due to a
wider opening is very small, and does not compensate for the
necessary increase in gross length of armature. The ventilating
plates usually consist of iron stampings similar in shape to the
armature stampings, but thicker. Radial spacers of no great
width, but of sufficient strength to resist crushing or bending,
are riveted to the flat plates; they are so spaced as to coincide
with the center of each tooth, and allow the air to pass outward
by providing a number of small openings on the cylindrical
surface of the armature.

Openings must also be provided between the shaft and the
inside bore of the armature through which the cool air may be
drawn to the radial ventilating ducts. The radial spacers on
the ventilating plates assist the passage of the air through the
ducts, their function being similar to that of the vanes in a
centrifugal fan. Apart from the ventilating ducts, the outer
cylindrical surface of the armature is effective in getting rid of a


large amount of heat, and the higher the peripheral velocity of
the armature, the better will be the cooling effect.

When forced ventilation is adopted, a fan or centrifugal blower
may be provided at one end of the armature. This may as-
sist the action of radial ventilating ducts, or it may draw air
through axial ducts. When axial air ducts are provided, the
radial ventilating spaces are omitted, and the gross length of the
armature may therefore be reduced. The ventilation is through
holes punched in the armature plates which, when assembled,
will provide a number of longitudinal openings running parallel
with the armature conductors. These openings may be circular
in section and should not be less than 1 in. in diameter, especially
when the axial length of the armature is great, because they will
otherwise offer too much resistance to the passage of the air,
and will also be liable to become stopped up with dirt.

One advantage of axial ducts which, however, can only be
used with forced ventilation is that the heat from the body
of the armature can travel more easily to the surface from which
the heat is carried away than in the case of radial ducts. When
the cooling is by radial ducts, the heat due to the hysteresis and
eddy-current losses in the core must travel not only through the
iron, which is a good heat conductor, but also through the paper
or other insulation between laminations, which is a poor conductor
of heat. As a rough approximation, it may be said that the
thermal conductivity of the assembled armature stampings is
fifty times greater in the direction parallel to the plane of the
laminations than in a direction perpendicular to this plane. For
this reason, radial vent ducts, to be effectual, must be provided
at frequent intervals. The thickness of any one block of stamp-
ings between radial vent ducts rarely exceeds 3 in.

The coefficients for use in calculating temperature rise are
based on data obtained from actual machines, and owing to
variations in design and proportions they are at the best un-
reliable. When forced ventilation is adopted whether with
radial or axial vent ducts it is possible to design the fans
or blowers to pass a given number of cubic feet of air per second,
and the quantity can readily be checked by tests on the finished
machine. The design of such blowers does not come within the
scope of this book, neither is it possible to discuss at length the
whole subject of ventilation and temperature rise. For a more
complete study of this problem, the reader is referred to other


sources of information. A very good treatment of the subject
will be found in Chaps. IX and Xof PROFESSOR MILES WALKER'S
recently published book on dynamo-electric machinery. 1

A good practical rule for estimating the quantity of air nec-
essary to carry away the heat when forced ventilation is used
is based on the fact that a flow of 1 cu. ft. of air per minute will
carry heat away at the rate of 0.536T watts, where T is the
number of degrees Centigrade by which the temperature of the
air has been increased while passing over the heated surfaces.
Thus, if the difference in temperature between the outgoing

Online LibraryAlfred StillPrinciples of electrical design; d. c. and a. c. generators → online text (page 9 of 30)