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5

Mechanical Design Considerations for High- and Low-Pressure Casings

Introduction

The turbine casing is essentially a cylindrical vessel and the main stationary portion of

each expansion section. This casing encloses the rotating elements of the unit and at the same

time locates the stationary blades, either directly, or through the location and support of an inner

casing, which itself carries the stationary blades and/or diaphragms. The principle components of

the casing are the shells, which provide the mechanical strength of the element and carry and

locate other elements such as packing heads, diaphragms, and the inner casing or blade carriers.

The casing is normally split along its horizontal joint at the centerline to facilitate

assembly and provide access to the rotor and internal stationary portions of the unit. The shell

halves are normally connected through a bolted flange at their horizontal joint and act to contain

the working fluid while maintaining it in intimate contact with the steam path blade elements.

Casings may also provide locations for internal packings or portions of the steam seal

system and could, if moisture is present in the steam, be equipped with internal moisture

collection and drainage systems. The high-pressure shells should also, in the case of minor

failures, be capable of containing missiles that are generated from the rotor.

Both the upper and lower portions of the casings can be arranged to provide connections

for welded pipe stubs. To these stubs are connected external pipes that allow steam to be

extracted for regenerative feed heating or other cycle or process uses. Such steam is extracted

from the main steam flow. The casing may also be penetrated by other pipes that are used to

introduce or extract steam for other parts of the cycle. It is normal for pipe connections to the

upper half to be connected through flanges or other device that allows their quick disassemble at

outages and then reconnection without the use of any form of heating or metal fusion techniques.

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Special provisions in the casing are necessary to admit the high-pressure, high-

temperature steam and make provision for the differential expansion that occurs between the

various portions of the shells. Such differential expansion occurs because of the different

temperatures or temperature gradient along the axial length of the casing and also because of the

different rates at which the various parts of the unit heat and cool with main steam temperature

changes. For double-shell construction, it is necessary for the main inlet pipes to pass through the

outer casing and introduce steam to the main steam inlet belt or nozzle box.

The high-pressure casings are normally supported at each end through arms that are

produced integral with and extend from the casing to pedestals that are located adjacent to, and

between, the casings or sections. Transverse and/or axial keys are used to maintain alignment of

the shells at these pedestals. Such keys have normally been hardened by nitriding and are located

on the bottom vertical centerline to ensure correct alignment is maintained at all loads and during

transient operating conditions.

Low-pressure casings are designed to contain the steam and to minimize the in leakage of

air when the exhaust pressure is sub-atmospheric. Because at exhaust from the turbine the

volumetric flow is large, it is normal for these low-pressures elements to be produced by

fabrication, and because such fabrications are not structurally strong, it becomes necessary to

support them for their entire perimeter at their horizontal joint or a similar location below this

joint.

Components Comprising the Turbine Casing

The turbine casings have a number of individual elements, which when assembled allow

the unit to operate safely and to achieve high levels of reliability and efficiency. A list of the

principle components follows.

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The shells. The shells are the main structural components that are produced by casting,

fabrication, or in some designs by a combination of both depending upon the experience and

preference of the designer.

The shaft-end packing head. The packing head is attached to the shells, and carries the

gland rings that are located where the rotor passes through the shells. These heads are designed

to carry gland rings that minimize the outward leakage of the steam or the inward leakage of air.

The inlet section. The inlet to the steam path must be designed to allow free access of the

inlet pipes, transport the steam to the nozzle box, and minimize leakage of steam at those

locations. The inlet is designed to permit movement between the inlet pipes and the main body of

the shells.

The explosion diaphragm on low-pressure sections. In the low-pressure shells there is

a need to provide for the rapid removal of steam from the internals of the casing in the event

there is a sudden and high rate of pressure increase due to some transient condition.

A diffuser section at exhaust from the last stage. In an effort to maximize the energy

extracted from the working fluid, the final rotating blade is arranged to exhaust into a diffuser

section normally produced as part of the casing fabrication.

Functions of the Shells or Casings

The casings are the main containment vessel of the turbine, which defines their major

function of containing the working fluid. These elements, therefore, surround the steam path and

can be made to perform several other secondary functions. These functions are dependent upon

the steam conditions within the steam path, which conditions influence the arrangement,

materials, and support for these casings.

It is necessary to consider casings in two separate categories, arranged most suitably by

the temperature of the steam they contain. The high-pressure/temperature elements are those that

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operate with steam temperatures above the range of about 700 degrees Fahrenheit (°F), and the

low-pressure/temperature elements are those that operate at temperatures below this value.

The high-pressure/high-temperature sections

The steam admitted to these high duty casings will have pressures up to 3500 pounds per

square inch absolute (psia) although units have been designed to 5000 psia. Temperatures are at

the 1000 to 1100°F value at the inlet and can be reheated to 1000 to 1050°F before readmission

to the intermediate pressure section. However, units have been designed to operate at higher

values of temperature.

Because of the initial steam conditions, the high-pressure/high-temperature casings are

subjected to high internal pressure, which produces significant tangential, longitudinal, and radial

stresses in the walls. These casings must therefore be designed so they are able to withstand

these conditions at normal operating conditions and during transients. It must also be recognized

that at the higher temperatures, the mechanical properties of the material from which the casing

is produced are lowered, which reduces the factors of safety of these major components.

The reheat casings are subject to lower steam pressures, but because of the increase in

specific volume of the steam at these lower pressures and the high reheat temperatures, these

casings can be subject to stresses of the same magnitude as exist in the high-pressure sections.

In addition to being a containment pressure vessel, the shells have certain secondary functions.

Fulfillment of these functions is important to the production of a successful design and is

necessary for the operation of the unit. The most important of these follow.

• The outer shells of high- and intermediate-pressure sections are part of the main

external structure. As such they must have sufficient strength they are able to transmit

the large differential expansion forces through the casing arms to slide the pedestals

on their sole plates. This they must do without any form of vertical or lateral

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distortion that would affect the alignment of the steam path. The sole plates must be

able to be moved by the casing in such a manner they will not cause deflection,

excessive distortion, or misalignment. Alignment of the turbine generator must be

maintained at all times, under all loads and variations of steam conditions, and in all

directions.

• The casings must also be able to carry the loads developed on the stationary blade

rows (individual blades and diaphragms) and inner casing due the pressure deferential

within the steam path.

• The shells should have sufficient strength and weight that the casing is able to resist,

without change of alignment, external forces and moments imposed on it by station

piping. (Tavernelli and Coffin 1961) Figure 5–1 shows how various forces may be

imposed on the casing by expanding piping thrusts that tend to lift the casing from its

foundations and could be sufficient to cause misalignment.

Fig. 5–1 The Piping Thrusts Developed on a Casing

• The shells must maintain the stationary elements they carry in correct axial and radial

alignment relative to the rotor. That is, concentricity must be maintained together

with axial alignment.

• The outer shells must be sufficiently rigid that they are able to transmit and withstand

external forces due to excessive vibrations, including earthquakes and other high

intensity natural phenomena. These phenomena, although rare and highly unlikely in

most North American installations, could have catastrophic consequences if their

severity were sufficient to cause sudden and excessive misalignment within a casing

with the rotor at operating speed.

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• The mass of the casings must be sufficient to make significant contributions toward

holding the unit firmly on its foundations and suppressing vibrations.

• In the event of a blade, wheel, or rotor failure, the inner and outer shells of the high-

and intermediate-pressure casings provide a strong containment vessel for the rotating

parts. These shells should be capable of absorbing high impact projectiles, thereby

minimizing the possibility of a projectile penetrating the casing and causing serious

injury to plant personnel.

• The outer shell provides a barrier by which heat is retained within the unit. This

barrier is reinforced by thermal lagging, which is attached to the outer surfaces of the

shell, and is the main barrier to radiant heat loss. The inner casing also provides a heat

barrier which reduces heat loss by minimizing temperatures on the inner surface of

the outer shell.

• Turbine shells are massive structures. They are thick sectioned, and due to their mass,

respond slowly to changes in steam temperature. This thermal inertia to the rate of

temperature change, gives rise to the need for special considerations of stationary to

rotating element clearances. The shell design must be adequate to accommodate this

thermal lethargy at all points of contact with potentially lower temperature elements,

such as valves, bearing housings, and the front standard.

• The thermal gradient developed in the casing walls will introduce thermal stresses

during operation. This is particularly so during temperature transients, when stresses

can be high.

The design of the shells should be such that the stresses induced provide a unit in which

the predicted life of the components is acceptable. To do this the material properties must be

carefully defined, and the design must eliminate, to the greatest extent possible, stress

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concentration regions. This will help minimize the possibility of thermal or low-cycle fatigue

cracking.

The high-pressure, high-temperature casings are normally cast components. However,

fabricated elements have been used in some nuclear applications where the initial or nuclear

boiler delivery pressures are not more than about 1000 psia. Such nuclear casings do, however,

have free moisture in them that introduces another type of problem.

The low-pressure/low-temperature sections

The low-pressure or exhaust sections of a turbine unit are normally designed to accept

steam at an inlet pressure of about 200 psia and a maximum temperature of about 700°F. This

maximum temperature is set more by the material of the rotor than of the casings.

The normal design practice is to make the total expansion ahead of the low-pressure

portion of the unit occur in one or more sections. At the lower pressure end of the high-pressure

casing, pressures may be at the 400 to 600psia level and temperatures in the 700 to 600°F range.

At exhaust from the intermediate or reheat section, the pressure will normally be in the range 70

to 200 psia and the temperature at the 550 to 700°F level.

The normal arrangement of the low-pressure expansion sections of a high-output unit is

to have multiple double-flow sections, with an inner and an outer casing in which the axial thrust

is canceled. In these designs, the casings at their inlet are subject to a pressure differential across

their walls equal to the differential between the inlet pressure and atmosphere. There and also

many units in service with three low-pressure expansions—one accepting one-third of the steam

exhausting from the reheat or intermediate pressure section, and the other two-thirds going to a

double-flow low-pressure section.

The low-pressure casings have many of the same functions and characteristics as the

high-pressure, high-temperature components. However, due to their physical size and the fact

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they are vessels required to maintain an internal pressure that is higher than atmospheric pressure

within the inner sections and a vacuum in their hood and between the inner an outer sections,

these requirements are modified. The basic functions of the low-pressure casing are:

• the outer shell must locate from the foundations and support the inner shell with

sufficient rigidity that it can maintain alignment of the steam path under all conditions

of transient load and steam conditions

• the inner casing must be able to carry and support the low-pressure diaphragms, to

maintain concentricity and axial alignment under all steam conditions and under both

steady state and transient loads.

• the casings must act as a transition and diverting structure to direct the steam

exhausting from the last stage blades to the condenser, minimizing the frictional loss

within the hood

• the low-pressure section casings must incorporate a seal system that limits the ingress

of air into the system and thereby help maintain vacuum integrity

• the casings must be sufficiently that robust they will not deflect by unacceptable

amounts due to vacuum pull during operation. Similarly the casing must be able to

resist vertical deflection due to heavy water loads in the condenser hot well.

• the casing, while mounted on the condenser with either rigid or flexible connections

and supported off the foundation, must have sufficient axial flexibility it is able to

accommodate temperature swings within the system and maintain alignment

• The casing must be designed with sufficient axial clearance to accommodate thermal

differential expansion at normal operating conditions and under short and long rotor

conditions

• The casings must be designed so steam extraction pockets can be used to remove

steam from the casings for regenerative feed heating

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The high and reheat casings can normally be expected to contain any blades that detach

as missiles. In addition these cast steel casings should also contain the rotor, although the unit

would be wrecked. Such high condition casings therefore act as a containment vessel or safety

barrier in the case of a significant accident or material rupture. The low-pressure casing may

contain the blades, although last stage blade elements can cause significant damage if they

detach from the rotor. If a rotor or wheel bursts in the low-pressure casing, it is most unlikely the

casing will be able to contain the missiles that are generated.

High-Pressure/High-temperature Casings

There are a number of casing configurations that fall within the category of high-

pressure/high-temperature application. These include the following.

High-pressure sections for fossil application. These units are normally subject to a

maximum cycle condition of 3500 psia and 1000°F. Although pressures up to 5000 psia and

1200°F have been used on advanced cycles. These casings are always built to the configuration

of an inner and an outer shell so that a pressure, and more importantly, a temperature gradient

can be established across both the inner and outer components.

Intermediate-pressure sections for fossil application. There are still in operation

turbine units that do not utilize reheat at exhaust from the high-pressure section. Therefore, there

are turbine casings that are intended to operate on steam having conditions equal to those

exhausting from the high-pressure section. Such units were used when the cost of fossil fuel was

inexpensive at the time the plant was built and are often used where a plant is located near the

fuel source. The intermediate pressure sections are probably the least stressed type, which can be

called high-pressure/high-temperature, and which will be encountered in modern power plants.

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Reheat pressure sections. It is normal in modern power plants to reheat the steam after it

has completed its initial expansion in the high-pressure section. Therefore, the steam entering the

intermediate pressure section casings has a pressure about 7 to 10% lower than the steam

exhausting from the high-pressure section, and the temperature of reheat is about 1000°F.

There are designs in which the high and reheat expansions are contained within a single

shell. In these designs the steam, after its initial (high-pressure) expansion, is returned to the

boiler reheat section, reheated, and returned to the same shell for a second expansion.

Second reheat sections. Some cycles are designed to utilize a second reheat section. In

this cycle the steam, after expanding in the first reheat section, is returned to the boiler where it is

given a second reheat and again returned to the turbine to continue its expansion in a second

lower pressure reheat section. Upon return from the boiler second reheat section, the steam has

had its temperature again raised to a value close to the initial temperature and is returned to this

second reheat section with a pressure reduced by 7 to 10% from that exhausting from the first

reheat section.

Those two turbine sections discussed previously are defined as the first and second reheat

sections and in certain applications are arranged for double flow.

High-pressure sections for nuclear application. With the advent of water-cooled

reactors producing low-quality steam, a high-pressure section was required that was capable of

handling large volumetric flows of steam that contained a small initial moisture content.

Therefore, as steam enters the turbine, it has an initial pressure of about 1000 psia and can have

an initial moisture content of 0.25%. In such casings, provision must be made to collect and

drain a considerable amount of water that will be deposited on the casings and other internal

parts of the unit as the pressure decays.

These turbine sections, in order to be able to accommodate the high volumetric flows

without exceeding axial velocity limitations for efficient expansion of the steam, have tended to

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be used at 1800 revolutions per minute (rpm) for 60 Hertz (Hz) applications. While the majority

of 50 Hz application is as 1500 rpm, there are some 3000 rpm applications at lower ratings. The

need to go to half speed units has caused an increase in rotor and casing diameters to maintain an

acceptable velocity ratio _. This has tended to increase the stress levels in the casings because of

the larger diameter required of the casings.

The casings for fossil application discussed here normally contain a high speed

rotor—3000 or 3600 rpm—driving a two-pole generator. Because of its high speed of rotation

coupled with high operating temperatures, there are physical limitations to the diameter that can

be specified for the rotor. Currently, it is difficult to produce a rotor forging with suitable

material properties and capable of carrying the rotating blades much larger than 40 inches (in.).

Also the maximum length of blades must be limited because of the centrifugal loading.

Therefore, the maximum casing internal diameter would be limited to about 65 to 70 in.

In nuclear high-pressure sections and some fossil sections, particularly for cross

compound units, the sections can be arranged to drive half speed 1500 rpm or 1800 rpm four-

pole generators. Because of their lower speed, it is possible to increase the rotor diameter without

exceeding stress limitations in the rotor or blades. With this type of rotor, a limitation of

approximately 64 in., producing a total rotor diameter of about 95 in. exists.

These diametral limitations are for 60 Hz units. For 50 Hz units, the possible diameters

would be somewhat larger. However, the maximum diameter is often a function of rotor

manufacturing capability rather than stress levels. As manufacturing techniques improve, it is

possible larger diameter rotor forgings will be available and larger casings required.

Pressure Staging and Multiple Shells

The casings contain the high-pressure, high-temperature steam with a differential from

working condition to atmospheric. The duty on the individual casing shells is normally reduced

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by the use of a double casing construction. This form of construction provides for a

temperature/pressure barrier to be established with conditions from high-pressure inlet to high-

pressure exhaust across an inner shell,and then from high-pressure exhaust to atmospheric over

an outer.

Shown as Figure 5–2 is a two-casing arrangement, in which the individual diaphragms, or

stationary blades, are located and carried in the inner shell. This inner shell is then supported

from flanges machined into the outer surface of the inner shell. These locate in special locating

grooves machined into the inner surface of the outer shell. It can also be seen from Figure 5–2

that there is a constant pressure and thermal gradient across the outer shell, and the inner shell is

subjected to a gradient dependent upon the differences between the stage conditions and the

high-pressure exhaust surrounding the inner shell.

Fig. 5–2 A Double Casing Unit With the Diaphragms Carried in Inner Casings or Blade

Carriers

Figure 5–3 is a casing design with the high- and reheat-pressure sections are contained

within a single casing. With this arrangement the high-pressure expansion has an inner casing to

carry the diaphragms, and the outer casing is subject to the same pressure gradients as seen in the

casing design of Figure 5–2. After reheating, the steam is returned to the reheat section, which is

a single casing design, with the diaphragms carried in grooves machined into the inner surface of

the shell. Therefore, this shell is subject to a decreasing gradient along its length from stage

conditions to atmospheric.

Fig. 5–3 A Combined High-Pressure and Reheat Section

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For the lower-pressure/high-temperature shells, such as used for intermediate or reheat

sections, the major concern is with thermal gradients. In this type of section it is possible to make

the casing walls thinner and more flexible. A casing of this design is shown in Figure 5–4.

(Hummer and Drahy 1964) With this type of design, steam is introduced from the inlet pipe into

the inner casing. This unit has seals at a to prevent the excess leakage of steam while allowing

for the expansion and contraction of the inlet pipes. This allowance is provided to accommodate

pipe movement during start-up and shutdown, or whenever the inlet pipes will heat and cool

much faster than the surrounding casing.

Fig. 5–4 An Intermediate (Reheat) Section, With Inner Walls and Extraction Pockets for

Pressurizing the Outer Casing

The steam enters the nozzle box then expands through the steam path. At completion of

its expansion, steam at the high-pressure exhaust condition surrounds the accessible portion of

the inner casing, which is then subject to pressure and temperature gradients corresponding to the

difference between individual stage and section exhaust conditions. The outer shell is subject to

pressure and temperature differentials equal to the high-pressure exhaust conditions and local

ambient. Had this casing been of single-shell construction, the single outer casing would have

been subject to the total differential between stage and ambient conditions.

In this type of design, it can be seen that diaphragms are supported and carried in inner

casing rings with each supporting a number of stages. These are also termed blade carriers.

These diaphragm groups provide for access regions where steam can be removed from the unit

for regenerative feed heating. In these extraction belts or pockets, the steam exists at the stage

discharge conditions from the upstream carrier, making the outer casing inner surface conditions

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equivalent to the extraction conditions. This reduces the temperature and pressure gradient across

the casing walls.

An alternate arrangement of the high-pressure casing is to eliminate the inner shell after

the high-pressure expansion is partially complete. When this is done the direction of the flow is

reversed, led to the other side of the nozzle box where the expansion is completed. Shown as

Figure 1–38 of chapter 1 is such a design with the flow reversed after eight stages to flow

through a final three before being returned to the boiler reheater section. With this design, the

steam path is split into two portions, an upper pressure portion and a lower pressure portion. The

arrangement of the shaft-end seals is the same except the pressure range across them will differ.

The only significant difference in such a design is that the outer casing will be subject to

a higher pressure and temperature differential over the first portion of the expansion. The second

or reversed portion has eliminated the inner portion of the casing, and the diaphragms are carried

by the single casing.

The reversal point—end of expansion portion—is selected based on three considerations.

These are:

1. the need to extract steam from the section for regenerative feed heating

2. the need to lower the temperature and pressure gradients across the individual casing

portions

3. the adjustment of the axial thrust developed in the two blade portions. These two

portions of thrusts are opposed, and will affect the thrust which needs to be carried by

the thrust block.

Reversal point selection affects shell pressure, temperature, and axial thrust. These effects

are best reviewed from the high-pressure expansion line of Figure 5–5 that shows that steam

enters the section at conditions Pin and Tin. The individual stage points of the high-pressure

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section are established along this expansion line, and possible stage end points are shown as a0

… a3 providing for optimum velocity ratios _ of the stages.

Fig. 5–5 The Reversal Effect and the Selection of Pressure and Temperature at the

Reversal Point

The design evaluation will consider the impact of different reversal points and the effect

these will have on the turbine and cycle efficiency. Normally, the controlling consideration is

achieving an extraction point for regenerative feed heating as this extraction will normally be to

the top heater and will therefore set the final temperature of the feed water being returned to the

boiler. This temperature is fundamental in establishing the heat rate of the total installation.

There is some small degree of flexibility in selecting the pressure and temperature at the

reversal point. This flexibility is achieved by selection of stage diameters that will modify the

velocity ratio _ and the energy distribution across the individual stages above the reversal point.

From the expansion line of Figure 5–5, the possible reversal points are shown as a1, a2 or

a3, with a0 being the inlet to the first of the three alternates being considered. These alternate

stage points will influence both the turbine and cycle. From considerations of the expansion line

alternates, the effect on the unit can be seen. In Figure 5–5b, the three stages are in series, with

their individual thrusts Tn acting in the same direction. In Figure 5–5c, the last of these three

stages, the steam flow direction has been reversed. Therefore, the steam will reverse at exhaust

from the second stage, and enter the third in the opposite direction changing the thrust by an

amount 2xT3. Also the temperature at the reversal point will increase from To3 to To2.

This change will also increase the steam condition surrounding the inner casing and

modify the thermal gradients across both inner and outer casings.

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The modern design many manufacturers utilize contains a nozzle box as shown in Figure

5–2 and in Figure 1–38 of chapter 1. These nozzle boxes are self-contained vessels, located

within but forming part of the inner casing. These boxes are normally produced by forging and

are designed to distribute the steam around a portion of the inlet annulus and discharge it through

the first stage nozzle plates as seen in chapter 6. In the case of the nozzle box, the highest steam

conditions sensed by the casing are those of the steam discharging from these first stage nozzles.

The total casing arrangement in a self-contained nozzle box is essentially that of a triple-

shell construction. Nozzle boxes are now used in practically all designs with an initial pressure

above 2000 psia and temperatures above 900°F. Depending upon the duty intended for the unit

and the system into which it will be electrically connected, the first stage nozzles may be

grouped in the following manner.

Segmental or nozzle control. If four or more physically separated inlet segments, as

shown in Figure 5–6a, are used the unit is termed nozzle controlled. With this design, admission

to each segment is controlled by a separate valve. The valves are each arranged to open or close

sequentially as unit output demand changes. The nozzle segments cover the complete 360° inlet

or whatever portion is required to access sufficient steam to the unit. With this design, there is a

small portion of inactive arc at the tangential transition from one nozzle segment to another.

Fig. 5–6 Alternate Methods for the Admission of Steam to the First Stage of a High-

Pressure Section

In this design, the valve opening sequence is V1, V2, V3, and finally V4. As each valve

opens sequentially, the active arc grows in tangential or chord length dependent upon the load

demand on the unit.

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Two 180° arcs. A similar design employs two 180° segments, Figure 5–6b, with the joint

between the inlet arcs at the horizontal joint. In this arrangement, the inlet arcs may be fed by

one- or two-control valve arrangements. Again there are small inactive arcs, which in this design

are located at the horizontal joints.

As with the design shown in Figure 5–6a, steam is admitted to independent arcs, each

covering a nominal 180° of the tangential position. Up to 50% load steam is admitted to the top

half only with valves V1 and V2 open. Past 50% the other valves open to full load.

Full arc admission. One 360° segment or inlet arc is seen in Figure 5–6c. This

arrangement is similar to Figure 5–6b except there is a flow connection from the upper to the

lower chambers. This flow connection may be in the steam chest downstream of the control

valves but is more commonly made in a header adjacent to the control valves. Steam flow to this

common chamber is controlled by valves that admit steam to the entire inlet arc.

With this design the valves will open sequentially in response to load demands, but each

of the valves V1, V2, V3, and V4 feeds the complete 360° arc. This is termed throttle control.

There will normally be a small inactive arc at the horizontal joint. However, the effect of this on

the stimulus produced can be reduced by careful design of the joint partitions.

The Low-Pressure Casings

The term low-pressure/low temperature, in terms of turbine section arrangement is

applied to those expansion that accept incoming steam from a higher pressure section, and allow

it to expand to exhaust or condenser pressure. The casings that enclose this energy level steam

path tend, in modern units to be a separate, often double-flow section.

However, in many older and lower rating units without steam reheat, the lower steam

condition expansion occurred in a casing that was integral with the inlet or higher condition

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casing at the inlet portion of the expansion. Often these casings are built using a cast high-

pressure section with a fabricated low-pressure section.

The older design single-casing units were produced by both casting and fabrication. In

many designs, the high-pressure section was produced from a steel casting and the lower

pressure portion was a fabricated structure bolted at a vertical joint to the high-pressure section.

This joint will often have a seal weld around its outer diameter to prevent the flow of air into or

steam out of the steam path.

The low-pressure casing is designed to accept steam from the exhaust of the expansion

immediately above it in terms of system pressure and temperature. The conditions of the steam

admitted to low-pressure casing are typically as follows.

In a fossil cycle. In these cycles the steam derives from the high, intermediate, or reheat

sections. Such steam is normally superheated. Its pressure is generally in the range of 70 to 200

psia. The initial temperature can be as high as 800°F. However, there are often limits placed on

this temperature not by considerations of the casing but rather by the operating temperature the

low-pressure rotor material can tolerate.

In a water-cooled nuclear cycle. In these cycles, the steam is admitted from the

intermediate system of the unit. Such an intermediate system will comprise a moisture separator

and possibly a reheater. Therefore, the steam conditions are typically in the range 70 to 250 psia,

and the temperature in the non-reheat cycle is at the saturation temperature corresponding to the

steam inlet pressure. In the case of the nuclear reheat cycle, steam is raised to a temperature less

than the initial cycle steam pressure saturation temperature by an amount equal to the terminal

temperature difference of the live steam reheater.

The steam at entry to the nuclear non-reheat, low-pressure section can contain moisture,

and the quantity is a function of the effectiveness of the moisture separator. For this reason, it is

probable the low-pressure, non-reheat cycle will have moisture present throughout the

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expansion, and the casing must be designed to accommodate this moisture existing at high-

pressure levels and possibly having high velocities.

In an effort to maximize cycle efficiency and extract as much energy from the expanding

steam as effectively as possible, the exhaust pressure from the low-pressure section is passed to a

condenser that produces sub-atmospheric pressures in the low-pressure exhaust hood. The

condenser is normally optimized, designed, and selected to produce an exhaust pressure between

0.5 and 6.0 in. of mercury absolute (Hga) at all loads and with all cooling water temperatures.

As the exhaust pressure decreases there is an increase in the volumetric flow in the

discharge section of the L-0 blade system and casing. To minimize the frictional losses

associated with the resulting high-velocity flow of steam in the exhaust, the casing is normally

mounted directly above or adjacent and connected to the condensers. These exhaust casings can

also contain deflector plates designed to direct the steam into the condenser and distribute the

flow as evenly as possible over the entire flow down area.

The large volumetric flows associated with large modern units often requires multi-flow

exhausts be used so sufficient blade annulus area is available and the steam exhaust velocity is

limited to acceptable values.

Shown in Figure 5–7 is a double-flow low-pressure section with a monoblock rotor. From

this figure, it can be seen this section comprises a double-flow casing with five rows in each

flow. Both the inner and outer sections are fabricated. In this design the inner section is designed

to carry and support blade rings or diaphragms that are produced by the methods described in

chapter 6. The pockets used for the extraction of feed heating steam can also be seen.

Fig. 5–7 A Double-flow Low-Pressure Section

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The inner shell is designed to locate the stationary blade rows and hold them in a correct

spatial position relative to the rotating blade rows. The stationary blade carrier elements or rings

are produced as outer diaphragm webs that carry one or more stationary blade rows. These outer

blade-ring carriers or webs locate directly in grooves machined into the inner casing fabrication

and permit adjustment within the inner casing to achieve optimum steam path alignment. These

fabrications also allow for space to remove steam for regenerative feed heating.

In the upper half of the low-pressure casings there are pressure relief or explosion

diaphragms. These diaphragms are designed to rupture and relieve any pressure that exceeds

atmospheric. Therefore, if for some operational or other reason vacuum is lost and the pressure

inside the low-pressure hood increases to a value above an acceptable limit, then the reversal of

pressure will deflect the diaphragms out and cause them to rupture. Rupture of the explosion

diaphragms will release the inner pressure of the casing, allowing the steam to escape from the

unit into the power-house or atmosphere in the case of an outside unit. The diaphragm rupture

pressure is normally between 15 and 30 psia. Rupture of these diaphragms will automatically

shut down the unit. If pressure were allowed to build up in the exhaust hood, levels of pressure

and temperature would increase to levels that would destroy the blade system.

Low-Pressure Casing Arrangement

The large number of multiple exhausts required for modern condensing units is

conveniently achieved by arranging for two or more double-flow sections in parallel. To achieve

a suitable temperature increase rate in the feed heating system and because of the large energy

range in low-pressure sections, three or four extraction of steam for regenerative feed heating are

normally required from the low-pressure expansions. Such an extraction requirement means that

the low-pressure hoods be produced so steam can be removed at a number of stage points in each

5-21

expansion. This requirement can complicate the general arrangement, and for some designs,

demands steam path and hood variation from one flow to the other.

If steam flow quantities are such that a single-flow section does not provide sufficient

discharge area, it is normal to arrange the low-pressure portion of the unit to employ a single or

multiple double-flow low-pressure sections. It was common at one time to employ designs with

three exhaust flows, with the one single expansion connected to the intermediate or reheat

pressure section discharge. This concept is not used extensively in the majority of modern units,

as it is more cost effective to develop modular designs of double-flow units, with specific

arrangements for steam extraction. These modular low-pressure designs also permit a better

mechanical arrangement of the low-pressure sections.

It is, therefore, becoming less common to employ an arrangement of three exhaust flows.

There are, however, still in successful operation a number units in which a single-flow low-

pressure section is connected directly to the intermediate section. This intermediate-

pressure/low-pressure (IP/LP) section can then be used with a single double-flow section to

provide a three-flow arrangement. In the three-flow arrangement, the first stationary blade row of

the low-pressure sections is set so that steam admitted to each of the three flows is controlled so

that with possible different steam extractions patterns in each. The exhaust flow from each

expansion last-stage blade row is the same.

The pressure range across the low-pressure sections is small when compared to the high

and reheat sections at one-fifteenth to one tenth their range. However, the energy extracted from

the low-pressure section can produce an output comparable to the sum of the output from the

other two expansions. Because of its large physical size and the fact that the space between the

outer hood and inner casing is maintained at vacuum pressure, there is a large downward force

resulting from the pressure differential between the inner hood and atmospheric. This total

pressure is sufficient to deflect the total casing vertically downward.

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The low-pressure casing is, because of its size and the fact that it is not a massively rigid

structure like the cast high and intermediate or reheat sections, deflected downward. The extent

of this downward deflection is sensitive to the vacuum produced by the condenser. There can

also be a change in casing elevation as the level of water in the condenser hot well changes.

Older designs are still in use in which the total expansion from inlet condition to

condenser exhaust is achieved in a single casing. Shown as Figure 5–8 is the cross section of

such a unit, in which the casing is produced in sections that are bolted together to form a single

expansion. The low-pressure casing can be manufactured by either casting or fabrication. A seal

weld between the low-pressure and high-pressure casings may also be used. This casing is

designed to provide for the extraction of steam for regenerative feed heating and has a valve

chest produced integral with the high-pressure inlet.

Fig. 5–8 A Single Flow Unit With the Low-Pressure Casing Attached Directly to the

High-Pressure Section

Since the steam exhausting from the low-pressure section flows to the condenser, it is

convenient and economical to mount the low-pressure section above and connected directly to it.

While flexible connections exist, it is also convenient to weld the lower half casing to the

condenser shell to form a continuous structure.

The bearings supporting the low-pressure rotor can be constructed and supported in one

of two ways. These bearings are produced either with the bearing shells as an integral part of the

low-pressure fabrication or they are mounted external to the casing supported off the foundation.

There are two aspects of these two possible design alternates that should be considered and

evaluated.

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1. When mounted from a pedestal on the foundation, the rotor elevation does not

change with vacuum or condenser hot well water quantity. It then becomes relatively

easy to predict the deflected shape and alignment requirements of the rotating

portion of the unit. However, because the casing will deflect downward under these

influences and must carry the stationary portion of the steam path, including the

sealing arrangement at the shaft end and diaphragms, there could be a need to

increase the radial clearance of the sealing systems in the low-pressure section to

allow for the difference in vertical deflection between the two sets of steam path

components.

2. When supported from the low-pressure fabrication, the radial seals at both the shaft-

end positions and the stationary blades can be maintained at or near optimum values

because the rotor will rise and fall with the bearings. This will minimizes leakage

losses. However, because the rotor will rise and fall as the vacuum changes, the

designer must have data on predicted deflection amounts to be able to establish the

normal running deflected form of the rotor.

With the bearings located within the exhaust hood, the rotor will have a shorter span,

limiting the bending stress induced in it.

Low-Pressure Casing Structures

The physical size of many low-pressure casings, particularly for 1500 and 1800 rpm

applications, are so large the casing must be constructed in several sections using vertical joints

in addition to the necessary horizontal joint split.

The two-casing (inner and outer) design of units consists of several portions, and these

should be considered separately because there are significant differences between them. These

casing segments are described next.

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The outer upper shell

For the small exhaust-stage blade designs, it is possible to produce the entire upper

fabrication as a single structure. For large exhaust blade systems, the upper shell may consist of

two or more fabrications with these sections joined by bolted connection at the centerline. It is

necessary to break the structure into sections because of the shipping and handling restrictions.

There can also, in the largest fabrication, be limitations imposed by the size of machine

tools required to produce the components and the furnace size needed to complete any stress

relief requirements after welding.

The main structural components of an upper outer casing are shown in Figure 5–9. The

basic shell consists of a wrapper plate that provides the upper outer casing, and there may be

connections to this wrapper from the crossover pipe for steam admission. There will, in addition,

normally be provision for explosion diaphragms. The end walls are normally flat and must

provide sufficient distance from the exhaust or discharge line of the blades to the end walls. If

this space is not sufficient, there will be an unacceptable loss of the steam kinetic energy upon

impact with the walls causing an energy loss within the hood (see chapter 3). The wrapper will

require the use of reinforcing ribs and struts within the hood to provide strength against both

distortion and the downward atmospheric deflection.

Fig. 5–9 An Upper Outer Fabricated Casing

Design considerations for the upper hood require the wrapper and end walls be

sufficiently thick to resist deflection and distortion and be suitable for the vacuum pull. The outer

hood should also be designed so the bearings and steam seal components can be accessed

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without the need to remove the outer casing. There will also be provision for access ports, so the

unit can be entered without the removal of the hood.

It is a normal design process to make a vibration analysis of the exhaust hood and then

use the reinforcing ribs and struts to de-tune the fabrication away from coincidence with any

natural frequencies developed within the structure.

The outer lower shell

The outer lower shell is the primary support structure carrying the low-pressure turbine

section. It must be capable of withstanding the vacuum load on both the side and end walls and

the vertical downward thrust transmitted to it by the upper half casing though the horizontal

joint. This structure must, if the low-pressure bearing is an integral part of the fabrication, carry

the bearings and support the weight of the rotor. The total downward thrust due to vacuum load

and weight must be carried through the casing while maintaining adequate bearing alignment.

The total low-pressure load is transmitted from the casing to the foundation by the support

brackets located at the sides and possibly the ends of the unit. The arrangement of a typical lower

half fabrication (half section) is shown in Figure 5–10. In Figure 5–11 is shown the lower half

casings of both the inner and outer portion with the double-flow rotor supported from the

bearings in the lower half.

Fig. 5–10 A Half Portion of a Lower Outer Casing

Fig. 5–11 An Open Low-Pressure Section Showing the Horizontal Joints of the Outer and

Inner Low-Pressure Casings

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The outer lower shell is fabricated from carbon steel plates, and the shell is given rigidity

by the use of internal struts. These can be seen in Figure 5–12. A primary design consideration of

the lower half outer shell is its need to support and provide location to the inner casing and

possibly bearing cones. It is also important that it can maintain radial and axial alignment during

both normal and transient operation.

Fig. 5–12 A Lower Half Inner Casing Seen From Above the Horizontal Joint

The inner casing

The inner casing carries and supports the low-pressure section stationary blades and/or

diaphragms. These casings, in a double-flow configuration, contain at their center an inlet bowl

that accepts the incoming steam from the crossover/around pipes and directs it into the first stage

stationary blade row around the complete 360° flow annulus. The inner casing also contains

extraction steam belts that collect the feed heating steam from the main steam flow required for

regenerative feed heating. These belts extend around the complete blade outer circumference and

are connected to a pipe transporting the steam to the heaters.

In some designs, the lowest pressure heaters are located within the condenser body, but

the steam must still be transported from the extraction belt to the heater shell. Because the low-

pressure section will have moisture in several stages, there is also provision made in the inner

casing to collect centrifuged moisture or to locate the diaphragms that have provision for this

moisture collection. In this case, the low-pressure casing provides the drains that remove the

collected moisture. The lower half of an inner casing is shown in Figure 5–12, where the

fabricated arrangement can be seen.

This inner casing is normally a separate structure supported from the outer casing.

However, for older smaller rated units with a lower inlet temperature, it is possible to

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manufacture the inner shell as an integral part of the outer. But for higher inlet temperatures, it is

necessary to manufacture the inner shell as a separate structure to accommodate the excessive

thermal gradients and differentials that can develop across the walls. In these older, lower rated

units with a lower inlet temperature, the single casing fulfills the requirements of the inner

casing. Also, it is connected directly to the condenser and therefore subject to the transient

thermal and load conditions normally experienced by the outer. These casings can be subjected

to high loading but the designer will allow sufficient margin that stress levels are well within

acceptable limits.

The inner shells are essentially open-ended, cylindrical pressure vessels with admitted

steam expanding axially in both directions in the double-flow configuration. In double-flow

designs, steam is admitted into an inner cylindrical annulus where it divides to flow axially out

through the steam path and exhausting to the condenser. Because the two flows are essentially

symmetrical, the axial thrust developed on the casing is also symmetrical and balanced. The

thrust developed in the tangential direction is in the same direction on both flows and therefore

additive. The casing must be keyed at its connection points to the other portions of the

foundation to ensure these thrusts are constrained. Relative to the weight of the structure these

thrusts are small, but there is normally some provision for containing them, particularly within

the individual blade rows.

The steam is admitted to the double-flow casing through one or two openings on top of,

at the bottom, or on the sides of the casing. The numbers and locations of these openings are

determined by the steam volumetric flow rate and conditions. The number of crossover/around

pipes, and their sizing is chosen so the mean steam velocity in the pipes and inlet annulus is not

greatly in excess of 150 feet per second (ft/sec).

Steam for feed heating is extracted from the inner casing at points immediately after the

rotating blade row. A circumferential opening into which the steam can flow is arranged around

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the shell periphery. The opening of the circumferential annulus or belt and the extraction pipes

are sized so the steam velocity will not exceed about 150 ft/sec. A typical extraction arrangement

is shown in Figure 5–13. In this figure can be seen an arrangement incorporating a water catcher

belt that is arranged and positioned to collect and drain moisture carried in with the steam,

centrifuged into the belt from the rotating blades, or carried into it from the outer flow walls of

the casing.

Fig. 5–13 Details of the Fabricated Structure of Wrappers and Carrier Rings Required to

Achieve a Satisfactory Structure

The inner shell is normally surrounded on its outer surface by wet steam with a

temperature corresponding to the saturation temperature of the condenser pressure or exhaust

steam. The inlet temperature to the inlet bowl can be as high as 800°F although a more normal

value is 700°F. Therefore, it is clear there can be relatively large thermal differentials developed

across the inner casing at some locations. Many manufacturers elect to design their inner casing

with a heat shield surrounding the inner section to minimize this thermal gradient effect.

The circumferential bowl at inlet to the double-flow low-pressure section is located in the

center section of the inner shell, and the steam extraction pockets are spaced axially along the

length of the fabrication with each of these succeeding pockets at a lower temperature reducing

toward the exhaust. Depending upon the extraction points within the expansion, there can be

temperature differentials across the separating walls as high as 350°F. However, the actual

differentials across the walls may not be as high as the indicated steam temperature differentials

because of the moisture film coefficient on either side of the plate. The outer wrapper plate can

have the inner surface exposed to 800°F steam adjacent to the crossover bowl, and two inches

away in the axial direction the inner surface might be exposed to steam at 450°F.

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The outer surface at this same location is exposed to the wet, cool condensing steam if no

thermal barrier is used. These multi-directional thermal gradients can induce extremely high

stresses and possibly cause casing distortions. If the stresses induced by these thermal gradients

are in excess of the yield strength of the material, then the distortions of the inner casing could be

permanent. Such distortions of the inner casing could result in clearance rubs and possibly

broken welds on reinforcing ribs and struts and therefore leakage at the various steam tight joint

faces. Should these stresses induce ruptures in the joining welds within the casing these can be

extremely difficult to access for weld repair.

Cast Low-Pressure Sections

While the majority of low-pressure casings are produced as fabrications, there are a

number of manufacturers that find that casting is still suitable because it is economical, reliable,

and capable of producing an effective product. The material used can be either cast iron or steel.

The principal material is cast iron. There are two forms of iron in use—graphite iron and the

spheroid graphite type. Cast iron is a material that is very suitable for casting, and it produces a

good quality form. Unfortunately the simple graphite cast iron cannot be easily upgraded if

defects are found. However, the spheroid graphite is readily welded and is therefore a suitable

material. Cast steel can be easily upgraded. Many modern two-casing low-pressure designs will

employ fabrication for the outer casing and casting for the inner casings.

Shown as Figure 5–14 is a large casting for an inner casing being turned after completion

of machining. The outer section into which this casing would be mounted would be produced by

fabrication. Shown as Figure 5–15 is the cast outer portion of a smaller unit.

Fig. 5–14 A Cast Inner Low-Pressure Casing Being Moved After Machining

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Fig. 5–15 The Cast Iron Exhaust Hood and Low-Pressure Casing for a Small Output Unit

The material specifications

When steel is used for the production of an inner casing, the material requirements for

these elements are generally not as stringent as for the higher steam conditions. A typical

chemical composition is shown in Table 5–1.

Table 5–1 Typical Composition of Low-pressure Turbine

In addition to these elements, a minimal amount of aluminum will be permitted for

deoxidation. Some manufacturers will also specify a small level of copper (0.30 to 0.60%) to

help combat and minimize the effects of washing erosion. The physical properties of this low

carbon steel are shown in Table 5–2. These properties are established from test coupons cast

integrally with the main casting.

Table 5–2 The Mechanical Properties of Low-pressure Turbine

The procedures for producing patterns molds and cores for these castings are identical to

those used for the high-temperature, high-pressure elements discussed in previously in this

chapter. With this type of casting, internal chills are not used and external chills are used only to

help achieve a logical solidification pattern and material structure. At completion of cooling, the

casting is shaken out from the mold, and the feeder heads are removed before the casting has

cooled below 400°F. The feeder heads must be removed in such a manner the steel is not burned.

Before machining, the casting is given a visual inspection for major defects.

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After rough machining the casting is given a nondestructive examination by magnetic

particle methods in accessible areas and radiographic examination in any weld preparation

regions

Acceptance level of casting defects

It is necessary to have established acceptance standards available for any casting faults

that might be found. The following criteria are intended to provide guidance only. There may be

other standards established by individual manufacturers for their units based on their

requirements and experience.

Visual acceptance standards. Folds, cavities, and clustered porosity with a depth greater

than 5% of the wall thickness should be removed by grinding. If the depth of the resulting cavity

is less than 10% of the wall thickness and the locations of the cavity are not in a region subject to

high stress levels, then these can often be accepted. It is best if the cavity is acceptable to blend it

out at its edges. If the cavity is greater than 10% of the wall thickness, then it should be weld

rebuilt and the requirements of stress relief applied.

Magnetic particle acceptance standards. Acceptable linear indications for critical and

non-critical regions are shown in Table 5–3. Surface defects are not permitted at planned weld

positions or at defect excavations. Discontinuous linear indications are considered acceptable

where the separation between adjacent indications is at least four times the length of the larger of

the two indications.

Table 5–3 Acceptable Magnetic Particle Inspection

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If as a consequence of this magnetic particle examination unacceptable defects are found,

then they should be excavated and weld repaired. After weld repair the casting should be

subjected to a stress relief cycle.

Defects in machined areas. The casting must be free from sub-surface defects which

would be exposed on machining. Defects classified as being greater than ASME schedule 1

found by radiography in scheduled weld regions are not acceptable. Any such defects in this area

should be repaired by welding.

Welding repairs. If it becomes necessary to weld repair defects in the cast shells, the

faults must first be excavated by some suitable means such as grinding and/or chipping,

machining, or arc-flame gouging. In some areas it is necessary to grind smooth the excavations

before the repairs proceed. The normal method of weld repair is manual metal arc. It is also

necessary to preheat the casting before repairs begin and to maintain the preheat temperature

throughout the repair procedure. Preheat temperature is from 150 to 300°F. Depending on the

material and whether localized preheat is used, this must extent for a least 10 in. in all directions

surrounding the repair. As the filler material is laid in, it must be continually inspected to ensure

no cavities remain in regions where they could lead to cracking as the unit ages.

Heat treatment. When the casting requires heat treatment, it should be loaded into the

oven and heated at a temperature ramp rate that should not exceed 200 to 225°F/hour (hr). For

annealing, the temperature should be raised to about 1700°F and for stress relieving to 1100°F.

These actual temperatures depend on the material. Once the treatment temperature has been

achieved evenly throughout the oven, these temperatures should be maintained for a period of

one hour for every inch of thickness of the thickest wall in the casing, but not less than a

minimum period of 12 hours.

Machining. The large physical size and weight of these castings, particularly for the half

speed (1500 and 1800 rpm) units, makes their handling and turning a complex operation because

5-33

they are normally much larger than the high and reheat section elements. The machining process

is essentially the same as for the high-pressure elements, and final machining requires the halves

be firmly bolted together and preferably supported in the manner as they are to be installed in the

field. If this is not done, these casings will tend to deflect when installed in the outer casing, and

the steam path grooving or support surfaces will no longer be concentric because the casing will

have a different sag form.

The machining of cast casings is essentially the same as for the fabricated. Shown in

Figure 5–16 is a large low-pressure section set up for internal boring where the two halves are

firmly bolted and machined as a pair.

Fig. 5–16 The Final Machining of a Low-Pressure Inner Casing

Thermal Gradient and High-Pressure Shell Design

There is, because of the energy expenditure within the steam as it flows through the

steam path, a considerable thermal gradient along the axis of any casing. There is also a thermal

gradient through the thickness of the walls of the shell due to the differential temperature that

exists across them. Under normal operating conditions, the casing can adjust to and

accommodate these gradients, and the shells can continue to operate satisfactorily for many

years. However, during operation there are changes in the temperatures to which the various

components are exposed, dependent upon the condition causing the change and the rate at which

these changes occur.

The change of steam conditions with the greatest influence on the casing are those

changes that occur rapidly and cause an increases in the levels of stress developed in the casing

walls. These stresses can be sufficient to induce failure due to the phenomena of low cycle or

thermal fatigue. Such failures can occur after a few thousand or even a few hundred such cycles,

5-34

dependent upon the severity of the temperature change and the stress levels and the normally

severe concentration of stress that exists at these points. Figure 5–17 shows typical cracks in a

high-pressure unit due to the phenomena of low cycle fatigue. (D.P. Timo 1970) This crack

initiates at a sharp female corner where stress concentration is high.

Fig. 5–17 Portion of a High-Pressure Shell Showing the Circumferential Cracks Formed

in the Filet Radii Positions as a Consequence of Thermal Cycling

It is of interest to consider the magnitude of stress occurring across any component due to

temperature mismatch between an inner hot surface and an outer cooler surface. Consider an

element of shell wall shown in Figure 5–18 where the temperature on the inner hot surface is

shown as T1 and on the outer cooler surface as T2. In this wall, the temperature gradient or

mismatch is _T. Consider the mean gradient or change of temperature in three cases.

Fig. 5–18 Temperature Profiles Through the Walls of a Casing Under Various Heating

Cycles

Linear temperature degradation. Figure 5–18a shows a casing under normal operating

conditions with a gradient that is practically linear from T1 to T2. In this case a compressive

stress will exist between the hotter wall and the neutral axis and a tensile stress from the neutral

axis to the colder surface. From the zero stress of the neutral axis to the maximum stress in the

outer fibers of the wall material, there will be a local temperature gradient _T equal to 1/2(T1-

T2), and the stress will have a maximum value of fs.

fs = ∆T . µ . E

1 - S (5.1)

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where

fs = stress

_ = linear coefficient of thermal expansion

E = young modulus

S = poisson ratio

Parabolic degradation. Normally the temperature gradient will not be linear, but will

follow some other distribution such as the parabolic seen in Figure 5–18b. In this case, the

maximum local temperature gradient _T is 2/3(T1-T2).

A condition that can subject the casing and other portions of the unit to high thermal

stress and where manufacturer’s recommendations should be followed in detail, is control of the

temperature ramp rates at start-up and shutdown when correct procedures can be controlled and

followed.

Hyperbolic degradation. The most severe conditions however exist at start-up or

shutdown when hot steam is initially admitted to the unit, washing the cold metal inner surface

with hot steam. Under these conditions hot steam flows suddenly through the unit, washing the

cold surfaces. The temperature gradient _T then approximates T1-T2 as seen in Figure 5–18c and

results in a significantly higher stress level in the outer fibers of the wall material. These stresses

become particularly significant in any region where there is high stress concentration such as at

section changes or where there are small fillet radii.

Sudden temperature changes caused by load shedding or boiler excursions also introduce

this situation. Sudden temperature changes when the unit is hot are possibly more sever than at

start-up because the material is hotter and therefore has poorer mechanical properties.

To the greatest extent possible manufacturers will avoid fillet radii that are too small.

Unfortunately some design requirements demand these be present. Also some manufacturing

5-36

techniques require and produce such radii as a function of the technique itself. Cast surfaces,

particularly those internal to the steam inlet annulus that cannot be inspected visually and are

difficult to access, are prime candidates for producing regions where stress concentration can be

high.

The temperature gradient, and therefore the thermal stresses developed in portions of the

shell, can be limited by adjusting the rate at which the boiler conditions change or by adjusting

the turbine start-up rate. Unfortunately it is inevitable that during the life of the unit, there will be

some start-ups in which the thermal stress exceeds the yield strength of the material in some

portions of the shell. Similarly, there will be uncontrollable excursions where these stresses are

exceeded. Such situations are regrettable but can and must be accepted. The immediate and

cumulative effects of the excessive plastic strains induced must be understood and be readily

measurable.

In an effort to understand and limit this effect, many turbine builders have introduced

systems of measuring, recording, and aggregating the contribution of each start-up temperature

change or excursion, whether the induced thermal stress is of a high or low magnitude, toward

initiating a surface crack. The level of temperature mismatch between main steam and initial

metal temperature at start-up—temperature change—is converted to a low-cycle fatigue index

(LCFI). A typical curve of such an index is shown in Figure 5–19.

Fig. 5–19 LCFI as a Function of Temperature Changes

When the summation of all individual indices over the years of the unit’s operating life

reaches 100%, there is a possibility a surface crack will have initiated. This may not be harmful,

and it may require a further 100% aggregation for the crack to propagate to a significant depth

and even more operation before rupture would occur. If cracks are discovered early in their life,

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then they can normally be removed by grinding. However, the removal of material will not solve

the problem created by temperature transients. If the crack is in a position where grinding can be

undertaken, then grinding can slow the rate at which the crack will propagate, but will not

prevent its reoccurrence.

High thermal stresses and the resulting accumulation of high individual indices can be

prevented by ensuring the temperature of the steam washing the inner surface is only slightly in

excess of the temperature of the core of the metal. If possible, the mismatch temperature should

be limited to values between -50 and +100°F, although acceptable outer limits are -175 and

+270°F. Here, a minus sign (-) indicates the main steam is cooler than the internal metal

temperature and a plus sign (+) indicates the main steam is hotter than the metal. The actual

values of acceptable temperature differentials for any unit will depend upon various factors

including the thickness of the metal section and the thermal conductivity of the shell material.

These recommended temperature differentials are normally the limiting factor to unit

start-ups, and the average thermal gradient during operation should not exceed the recommended

if an acceptable life is to be expected from the equipment. There may be some parts of the unit

where steady-state temperature differences larger than this will occur during normal operation. In

such cases, parts will have been designed to accommodate this and be of suitable materials and

form to provide the degree of flexibility required to prevent excessive stress.

Estimating Low-Cycle Fatigue Life

The model just discussed for considering the production of thermal stresses and their

effect on casing life together with the introduction of low cycle fatigue cracks, although very

simple, provides a relatively simple tool for estimating casing life consumption due to start-up,

shutdown, and during operation transients when large temperature changes occur.

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The life expenditure that occurs during any unit transient is a function of three

factors—the magnitude of the thermal stress induced in the shell, the material properties of the

shell, and the environmental temperature at which the change has occurred.

Any component repeatedly subjected to stresses beyond the yield strain of the component

material at its operating temperature will develop cracks in a finite number of cycles. The

number of cycles required to initiate these cracks is a function of the stress level.

Turbine shells are a relatively complex form, and they contain regions where during

operation there is a considerable degree of stress concentration in the parts that are subject to

biaxial loading. Therefore, the calculation of actual stress levels by traditional methods is

difficult, although finite element methods have allowed a much better understanding of the loads

and stresses involved in casings. Because of these difficulties, designers find it is of considerable

advantage to calibrate experimental values of stress against calculated values, which are

normally determined by finite element methods.

Figure 5–20 shows a portion of an outer casing scale model equipped with strain gauges

used to predict actual values. These stresses are then compared with calculated values, which

permits experimental factors to be established which can be applied to other casings with similar

geometries to obtain an adequate degree of accuracy.

Fig. 5–20 Scale Model of an Outer Shell Instrumented With Strain Gauges to Help

Establish Stress Levels

In any repeated stress/strain situation in which the stresses developed within the

component are in excess of the yield stress in tension and compression, it is usual practice to plot

the total (elastic plus plastic) strain elastic plus static (EET) against cycles in determining the life

factor for the component. Two curves for a typical shell material are shown as Figure 5–21. The

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materials selected for any component have a considerable effect on the low-cycle fatigue. In

Figure 5–21curve A being a low strength relatively ductile alloy and curve B a stronger, less

ductile one.

Fig 5–21 Elastic + Plastic Strain as a Function of Stress Cycles for Two Different Shell

Materials

It has been shown that at room temperature, low-cycle fatigue life may be predicted by

the following expression. (Tavernelli and Coffin 1961)

∆εp = 0.5 Ln [100/(100 - %Ra)]

N (5.2)

where

__p = plastic strain range

%Ra = percent reduction in area measured in a tensile test specimen

N = the number of cycles to cracking

This expression is valid in the high strain range where the ratio of plastic to elastic strain

is high. In any casing form, it is normal for the designer to determine the number of cycles N to

initiate a crack. This number is factored into the total design considerations including the

selection of the material to be used.

Thermal Gradient in the Low-Pressure Inner Casing

In many low-pressure sections, problems are encountered due to thermal gradients

causing permanent distortion of the inner casing. It is normal for these gradients to occur in both

the axial and radial direction. Therefore, the distortion which results can occur in a complex form

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in the casing structure causing permanent set in both directions. In addition to the predictable

temperature differentials from the expansion and extraction of steam, it is known that the skin

surface temperature of the outer surface of the inner casing varies in an unpredictable manner

and is influenced by changes in load and varies from the upper to the lower halves at any

transverse section.

At exhaust from the last stage blade annulus, the steam is deflected to flow into the

condenser. However, there are spaces between the inner casing outer surface and the inner

surface of the top half outer casing. These spaces fill with flowing steam, which passes through

them to the condenser and the surfaces therefore attain steam temperature. This represents a

thermal gradient on the inner casing walls. However, many designs of inner casings are arranged

to include a thermal barrier attached to the outer surface of the inner casing. This barrier helps

ensure the temperature gradient across the wall is not as severe as that caused by the outer

surface of the inner casing attaining steam temperature.

There are within the exhaust hood factors that cause temperature variation and non-

symmetric flow. These include:

• unit load. As the unit load varies, so will the quantity of steam flowing through the

steam path, which will in turn modify the flow velocities and patterns through the

spaces between the hoods.

• exhaust pressure. As the exhaust pressure produced by the condenser changes, there

will be a change in the steam specific volume and the volumetric flow will change.

Also, as the condenser pressure changes so will the saturation temperature of the

steam that covers the metal surfaces. There will be changes in steam velocity and

temperature associated with condenser pressure changes.

• Rotational effect. While the flow pattern of the steam at exhaust from the last stage

blade annulus will be substantially axial, it will possibly have some tangential and

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radial component to its flow. These effects will produce a total flow distribution that

is different from side to side within the casing. This side-to-side flow difference will

influence the total flow patterns within the hood, which will be sensitive to any small

change in either steam pressure or quantity.

In designing a low-pressure section, sufficient flow area must be made available to the

exhausting steam to minimize the pressure drop from the blade annulus to the condenser. There

are two important considerations to this requirement.

1. The exhaust blades will discharge their flowing steam into a diffuser that is produced

as part of the low-pressure section fabrication. The diffuser form is selected to

minimize losses associated with removing the steam away from the exhaust plane and

not impede further flow from the blades.

The diffuser is normally constructed from rolled plate that is either welded or bolted

to the inner casing. The axial distance from the exhaust blade annulus to the casing

end wall is limited, and it is difficult to achieve a perfect arrangement within the axial

length available. However, designs can be provided that allow a diffuser section to be

used and can be accommodated within the available axial space to help minimize the

losses which occur.

2. The hood structure must turn and divert the steam, normally downward, to the

condenser, causing a minimal frictional loss within the hood. Hoods are designed so

strategically placed diverter plates can turn the steam in an effort to keep the flow

density at any point relatively constant, avoiding excessive velocities and minimizing

frictional losses.

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During operation, the inner casing must remain sufficiently rigid it is able to maintain

concentricity in the radial direction and retain axial alignment. It must do this and yet remain

sufficiently flexible it is able to respond to large temperature swings that occur within short

periods of time. These requirements of flexibility and rigidity are obviously contradictory.

However, it is important that alignment and rigidity requirements are addressed in the design

phase. If any forced compromise is required by one requirement, then it must be recognized by

and accounted for in defining the requirements of the other. These requirements can be

aggravated by any large temperature gradients, and the designer must anticipate the most severe

condition when defining the low-pressure sections.

There are various approaches that have been considered to solving the thermal gradient

problems encountered in designing and manufacturing the low-pressure casings. An attempt

could be made to reduce or eliminate the radial and axial gradients by insulating the various

members of the fabrication. Also, the radial gradient could be reduced by insulating the inner

surface of the wrapper plate. However, any insulation used would be exposed to wet steam, and

unless this insulation was impervious to water soak, it would immediately loose its insulation

properties on becoming wet. A ceramic insulation would overcome water soak problems but

would be unable to expand and contract adequately to accommodate casing movement.

To reduce or eliminate axial temperature gradients, the shell would have to be made of

several cylindrical sections to minimize conductive heat transfer. This would require making the

inlet bowl and each extraction belt a separate fabrication. This is obviously an expensive solution

since it would require a multiplicity of transverse flange faces and could present considerable

alignment problems. In addition, thermal cycling of the casing affects the bolting on flange faces,

and each separate fabrication would need to be supported individually from the lower half outer

casing.

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Since there is considerable difficulty in reducing or eliminating gradients, the normal

engineering approach has been to design the structures so they are able to accept the anticipated

gradients and not have stresses induced in them which exceed the yield strength of the material.

This is the design approach currently pursued by manufacturers. It has so far proven to be an

acceptable solution, but the costs of producing the casings are increased by the use of more

expensive material and thicker sections in some locations than are required from the simple

consideration of normal (non transient) gradients.

The schematic of an inner shell, shown as Figure 5–22, indicates the predicted steam

temperatures and pressures at various locations within an inner shell arranged for steam

extraction pockets. These conditions are consistent with normal operation and will change during

transient operation. In this type of design, the only connections from the inner support sections to

the cooler wrapper plate are relatively thin supporting ribs. These ribs are free to deflect and

move axially under the influence of both thermal growth and diaphragm thrust. This type of

design eliminates the compressive stresses that would be present in the ribs if these had been

massive structures and the extraction pockets or belts had not been circumferential, thus

permitting limited axial movement.

Fig. 5–22 Temperatures at Various Locations in a Fabricated Low-Pressure Casing

High-Pressure Turbine Shell Materials

Advancing steam conditions and increases in diameter, particularly for half-speed

machines, have required a continual improvement in both the composition and mechanical

properties of the material and the manufacturing techniques used to produce steam turbine casing

shells. This is particularly important when applied to high-temperature, high-pressure units.

Casings are produced from alloy steels, and the castings are carefully controlled both to ensure

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mechanical strength and freedom from casting defects which have the capability to compromise

the integrity of the shell.

For temperatures up to about 750°F, a material that is produced to American Society for

Testing and Materials (ASTM) A27 Grade 65-35 will normally be acceptable. The mechanical

and chemical specifications for this material may be modified by closer control of the chemical

constituents and the heat treatment undertaken. However, the material will generally meet the

overall requirements of this specification. For increased temperatures a more suitable

specification is one that accords closely with the requirements of ASTM A356 Group 8, and for

the highest steam temperatures up to about 1100°F the ASTM A356 Group 9 specification is

most suitable.

Typical chemical constituents of these materials are shown in Table 5–4 and the

minimum acceptable mechanical properties in Table 5–5. Turbine builders will modify these

basic requirements to suit their particular philosophies, applications, and design requirements.

This is acceptable and reflects the experience from many years of operation.

Table 5–4 Turbine Shell Castings Nominal Chemical Composition

Table 5–5 Turbine Shell Castings Minimum Mechanical Properties

In addition to sulfur and phosphorus, there are normally other trace elements such as

antimony, arsenic, and tin, present in the material. However, these amounts should be kept to a

minimum. Aluminum is generally not tolerated and there are two basic reasons for this. First,

aluminum has a greater affinity for the elements with which vanadium should form a compound

to increase the strength of the material. Therefore, if aluminum is present it will effectively

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reduce the mechanical strength of the casting. Second, aluminum present in a casting operating at

high-temperatures weakens the component’s resistance to creep stress.

Therefore, freedom from aluminum and general purity is essential because these castings

will need to be weld repaired and upgraded after casting and before final machining is

completed. Therefore, high levels of impurities will compromise the strength and integrity of the

shells. Because of the heat fusion processes and then the stresses that will develop in the shells,

the control of the chemical constituents is critical.

The addition of a small quantity of copper acts to reduce the incidence of washing and

wire drawing erosion that can affect a casing. Therefore, for nuclear applications a small amount

of copper is often specified.

The service to which casings are put requires close control of their chemical constituents,

to help ensure their long-term service will be at an acceptable level. This will particularly require

the specified chemical content plus any nickel and aluminum be controlled so as not to exceed a

specified carbon equivalent. For the lower temperature material, ASTM A27 is the controlling

formula:

C + 0.333 [Mn + Cr ] + 0.167 [ Si + Ni ] + 0.500 Mo = 0.750 (5.3)

For decreases of carbon content below the maximum specified, down to a specified

minimum, the manganese content may increase, although a maximum level is placed on this.

Again the turbine designer may require slightly different mechanical properties than

those contained in Table 5–5. These requirements will be detailed in the material specification.

On the basis of present operating experience, it is considered the mechanical properties

for high-pressure/high-temperature shells operating below the creep range are adequate if the

materials exhibit a low Charpy impact transition temperature. (Rogers and Brewer 1964–65)

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The complex forms required for shells, coupled with high initial steam conditions, make

castings the obvious manufacturing process for producing these components. Many of the highly

stressed regions in the casing are required to be thick and these occur in regions subject to rapid

temperature changes. The hardenability of the alloy used to produce these shells should be

adequate to obtain the correct properties throughout the section. Since both inlet and outlet

connections must be made to the casings, the materials must be weldable.

Faults are a common factor of casting and repairs must be made by the deposition of weld

material. To permit adequate wall thickness and the inevitable stress concentrations, high rupture

strength and rupture ductility are necessary. (Curran and Timo 1964)

Steel Plate for Low-Pressure Casings

The low-pressure outer hood and inner casings are normally produced by weldments or

steel fabrications. These parts are large and subject to considerable loading due to both the

dynamics of the unit and the downward pull of the condenser. The low-pressure casings must be

free to move when subject to axial thrusts and yet maintain alignment. For these reasons, the

materials from which the fabrications are produced must be stable, able to be welded, and

maintain their long life integrity.

The steel used for these hood fabrications is a low or intermediate strength carbon-silicon

type. This plate is normally manufactured by the open hearth, basic-oxygen, or electric furnace

process. Because of the large amounts of free water that can be present in the low-pressure

section of modern units, it is often necessary to use stainless steel inserts at certain locations

because of the tendency of the carbon steel of the casings to loose material to both moisture

impact and washing erosion.

The plate specifications

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The chemical composition of the carbon steel plate used to manufacture both outer and

inner casings is relatively simple, the constituents being dependent upon the loads the

components will carry and the plate thickness. The steel contains silicon generally in the range

0.15 to 0.30 percent, the exact content being consistent with obtaining the required mechanical

properties and limiting the carbon content to the lowest practical level. (ASME Specification

A284-70a)

Despite this specified composition, there is normally some variation of composition

throughout the thickness of the plate.(Rollason 1956) This variation occurs as a consequence of

the rolled in inclusion from the initial ingot. There are also some impurities that segregate and

concentrate at other locations and are given a linear location during the rolling process. These

cannot be avoided but can be allowed for in the design of the component. Normal chemical

composition and mechanical properties are shown in Tables 5–6 and 5–7.

Table 5–6 Chemical Composition of Weldable Quality Steel Plate for Low-Pressure

Fabrications

Table 5–7 Mechanical Properties of Weldable Quality Steel Plate for Low-Pressure

Fabrications

In addition to the mechanical properties listed in Table 5–7, another important ductility

characteristic of this material is the bend test. It is normal for samples of this material to be bent

through 180 degrees without producing evidence of any surface cracking. The bend radius is

established as a function of the plate thickness.

Plate faults

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Plates less than about four inches in thickness are normally delivered to the turbine

manufacturer in the as-rolled condition. Therefore, before being used in any fabrication, these

plates must be descaled by shot blasting or some similar method. After this descaling process

surface faults are occasionally visible.

Molten steel contains soluble and insoluble gases and metallic and non-metallic

inclusions. When molten steel is poured into an ingot mold and cools, solidification starts at the

mold/steel interface and continues towards the center of the ingot. Segregation of the impurities

that normally have a lower melting point and the entrapment of gasses tend to take place at the

center of the ingot. The manufacturer can produce a rimmed steel, a capped steel, a semi-killed

steel, a killed steel, or a vacuum deoxidized steel. The degree of segregation progresses from

pronounced in a rimmed steel to virtually none in a vacuum degassed steel.

Therefore, the production process has a considerable effect on the quality of the steel.

Porosity in the ingot also follows the same gradation, going from considerable in the rimmed

steel to none in the vacuum degassed product. However there is a cost premium associated with

each of these processes, with cost increasing with the steps taken to limit the quantity and form

of the impurities present in the final plate.

When the ingot has been poured, the plate is rolled and any inclusions are flattened and

spread out. Dispersion takes place, and depending upon the degree of rolling, the resulting

platelets or inclusions may or may not be significant. Gases leave voids in the ingot, and with

sufficient hot rolling, even deep-seated blow holes will weld up and disappear. If not, the plates

will exhibit some level of lamination.

The manufacturing processes, including the accompanying segregation during cooling,

gives rise to certain types of faults in the plates.

The major internal defects

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Laminations. These are continuous faults that normally occur at or near the center

thickness of the plate. These faults are due principally to gasses or pipe that will concentrate at

the upper axial center of the ingot and become rolled into the plate. Figure 5–23 shows a cross

section of a typical lamination in a plate. Laminations can also be caused by rollover during the

rolling process. This rollover occurs when there is a doubling of the section during rolling. This

results in multiple layers or a pinching of excess material that is squeezed over the section

entering the rolls. These multiple layers form of lamination may occur at any thickness position

of the plate.

Fig. 5–23 Laminations and Lamellar Tearing in a Plate Adjacent to a Weld

Rolled out inclusions. During the cooling process, the dispersed non-metallic inclusions

can become trapped in the ingot in isolation from other impurities or defects and these are

surrounded by the steel. Being non-metallic, these inclusions form a non-continuous area in the

form of planes of incomplete fusion in the plate as shown diagrammatically in Figure 5–24a.

Fig. 5–24 Rolled Out Inclusions Near the Center Section of a Plate a, and the Lamellar

Tearing Under the Action of a Force P

In isolation, or in small groups, these small laminar type inclusions are not necessarily

detrimental to the strength of the plates. However, if they occur in the vicinity of a weld, their

effect can become significant. Under the influence of a weld and the internal stresses this metal

fusion introduces, decohesion of laminar-type inclusions will take place when the plates are

stressed in a direction transverse to the plane of the inclusions. When such decohesions occur in

the platelets that are close together, the resulting phenomena is known as lamellar tearing. This

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tearing has a characteristic morphology being step-like in appearance but tending to follow

planes parallel to the rolled surface of the plates.

The planar areas of decohesion are interconnected by near vertical shear faces normal to

the plate surface as seen in Figure 5–24b. For lamellar tearing to occur, two conditions must exist

in the plate or welded joint. There must be sufficient through-thickness stresses and strains on the

joint, and there must be low through thickness ductility.

The through thickness stresses and strains in a low-pressure shell can result from thermal

distortion during welding and later in service due to thermal loading particularly during transient

operation when there is heating and quenching of the shells.

When lamellar tearing occurs, the initial fracture is usually at a large inclusion to metal

interface. Regions of decohesion then link with other regions by shear fracture until

progressively the final linking of the larger areas results in gross cracking.

Lamellar tearing can be detected by nondestructive examination (NDE) methods.

Radiography is difficult to apply due to unknown crack orientation and accessibility. However,

for sub-surface cracks they are detectable by straight beam or angle beam ultrasonics. Figure

5–25 shows diagrammatically three typical lamellar cracking profiles that can be present in the

region of a weld.

Fig. 5–25 Various Tear Patterns Adjacent to Weld Joints

The major surface defects

Surface flaws. If a surface flaw or defect does not decrease the net cross sectional area of

the plate or result in potential stress raising effects, it can be ignored. Surface defects are in

general the result of faults and irregularities introduced during the steel pouring and rolling

process. Although surface defects are generally not detrimental to the performance and

5-51

mechanical strength of the plate, care must be exercised when forming or bending to ensure they

are not present in regions where they will contribute to any stress raising situations.

Ingot cracks. These cracks are the result chiefly of high pouring temperatures (values

considerably in excess of the solidification temperature). This high-temperature pouring results

in an inhomogeneous crystalline structure on cooling which has zones of weakness extending

from the edge of the ingot in towards the center. On rolling, these planes of weakness are sheared

and are present in the final plate.

Scabs. Scabs are caused by metal splash against the mold during the pouring process.

These splashes onto the side of the mold cool rapidly and oxidize. When the ingot is rolled, these

surface inclusions appear as surface scabs. This defect can be minimized by bottom pouring,

mold coating, and by close adherence to pouring rates and procedures.

Seams. These are the longitudinal cracks or openings that appear on the surface of semi-

killed steel. These are the result of longitudinal or transverse cracks in the ingot that are

elongated during the rolling process.

Burned steel. This form of damage is the result of flame impingement on the surface of

the ingot, usually at the corners. This intense heating causes oxidation at the grain boundaries

and results in a rupturing or tearing during rolling.

Cinder patches. These patches are usually caused by non-metallic pickup from the

soaking pits.

Laps. These are the result of overfilling of the mold, which in turn causes fins or

projections that are turned down in subsequent rolls. Laps are often deep and the plate cannot be

salvaged.

Casing Weldment Considerations

5-52

Fabrication from plate is the principal method for producing low-pressure casing

components. The duty and stresses to which these components are subjected demands attention

be paid to the detail of the welds, which are required to ensure a high degree of mechanical

integrity. In designing a fabricated casing, there are certain details of the structure that must be

considered that are fundamental to its strength and therefore acceptance. In general, to achieve

mechanical strength, the welds required to produce a suitable shell require full penetration. This

means that access to both sides of many structural plates must be available and there must be

sufficient space to permit internal welding of the structure.

During the life of the unit there will be many occasions, including the manufacturing

phase and during maintenance outages, when the casing upper halves particularly will require

lifting. To facilitate this lifting, the casing must be fitted with lifting lugs to allow the component

to be raised and moved without causing any form of twisting or significant distortion that could

in any way be permanent or affect its ability to be reassembled and maintain alignment and

running clearances. These lifts must be so designed that they will not overload individual

structural welds to the extent there is a rupture in any joint. Many upper half low-pressure

sections carry the diaphragms assembled into them. Therefore, when this upper half is raised the

diaphragms will lift with the casing, the casing must therefore be of sufficient strength it is able

to carry and support these diaphragm halves.

It is in the interest of the designer to select weld configurations that minimize overhead

and vertical welding to the greatest extent possible. While these forms of construction cannot be

eliminated entirely the individual pieces of the fabrication should be designed to permit as much

downhand welding as possible. Such considerations are of particular importance for site work,

which must be accomplished as quickly as possible and then be available for evaluation by non-

destructive means.

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The steels used for the fabrication of low-pressure casings do not require a preheat prior

to welding. However, at completion of the total fabrication and before machining, it is a good

and recommended practice to place the entire structure in a furnace for stress relief. Typically a

large fabricated structure should have a maximum heating rate of about 150°F/hr and be heated

to a stress relief temperature of 1100 to 1200°F for a minimum period of six hours. At

completion of this stress relief cycle, the structure should be cooled at a rate of 500 to

600°F/hour. The fabrication should not be air quenched.

In supporting the structure in the furnace, care should be taken to ensure all parts reach

furnace temperature. Sufficient thermocouples must be placed on the fabrication at strategic

locations to ensure a uniformity of temperature is maintained. The stress relief period should

begin when the specified temperature has been reached by all parts of the structure. In the

furnace, the casing should be supported so there is no sagging due to its own weight, otherwise

permanent distortion can result.

To provide a permanent record of the stress relief process, the following documentation

should be prepared and made available, as requested, to the purchaser.

• A photograph of the furnace load identified by serial number, material specification,

or other information allowing immediate and unchallengeable identification.

• A thermocouple chart showing individual location and thermal history of the heat-

treating process.

Should any repair welding be necessary after stress relief and inspection, it is necessary

to consider repeating the stress relief process. This will require evaluation and will be dependent

upon the amount of repair required and its location. The need for this should be determined from

an evaluation and the perceived need in relation to the stress and duty of the repaired area.

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At completion of the fabrication process, and sometimes during the manufacturing phase, NDE

of the weld joints should be undertaken to ensure the joints meet the engineering definition and

has therefore the greatest probability of achieving a suitable level of mechanical strength.

The quality of the welds is in general established from both visual examination and crack

detection techniques. Structural welds must be completed by qualified welders and the work

completed in the correct sequence to approved procedures.

The most suitable method for examining structural welds for integrity is by ultrasonic

examination, although it is possible in some locations to use magnetic particle inspection (MPI).

These examinations should be made after stress relief, and if considered necessary, before heat

treatment. The purpose of this evaluation is to ensure further upgrading and repair will not be

required after the stress relief. Hydrostatic testing is often used with large fabrications and has

generally been found to be satisfactory in proving the design strength and structural integrity.

This test provides a means of checking for leaks.

Smaller fabrications use oil as the filling fluid and larger fabrications use water,

sometimes with dye or dioxide in solution. The test pressures are usually specified in the range

of 1.5 times the maximum pressure the casing will experience. Some manufacturers have used

low-pressure neon gas for this test.

If a fluid is to be used to conduct the hydrostatic test, it is important that the casing be

filled with the test fluid and then left for a sufficient period so the fluid can reach a temperature

close to the ambient condition. A satisfactory test cannot be conducted when the casing is full of

cold fluid. Under these circumstances there is a tendency for condensation to form on the casing

surface, and it is not possible to differentiate between leakage and condensation.

Casing Weldment Details

5-55

There are many factors involved in the selection of the geometry for any particular joint,

and it is impractical to make specific recommendations in a general discussion. The selection of

geometry and certain other factors are dependent upon an evaluation of the position, loading, and

accessibility of the joint to be made. The unit designer is not normally in a position to specify

one particular weld without reference to the manufacturing function of a supplier. This is

because weld geometry is a function of the welding process to be used, and there could be

several suited to any particular location within the total fabrication. It is normal for the welding

procedure to specify applicable welding codes, and those requirements must be satisfied. The

manufacturing engineer is primarily responsible for meeting code requirements.

The stresses developed in a turbine hood or casing structure are difficult to determine

with any degree of accuracy, and model testing is the most appropriate method available,

although this is being supplemented and perhaps to a degree superseded by finite element

analysis as experience accumulates. However, low-pressure sections are normally a standard

design and it is only necessary to establish stress patterns and magnitudes once for each design

for the most severe operating conditions. When this has been established, future units are built in

such a manner these maximum stresses can be tolerated and the unit continues to operate in a

satisfactory manner.

In defining weld requirements it must be recognized that over-welding can be as serious a

fault as under-welding, because over-welding can induce residual stress capable of causing

casing distortion. This can be difficult to correct. In selecting the weld for any location within the

unit, the following three considerations need to be addressed.

1. If the weld is to provide a path for the transfer of forces, a welded design is justified,

and the calculations necessary to determine stress levels and suitable weld sizes and

geometries are mandatory.

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2. If the weld is simply to locate or hold parts together, full-length welds are invariably

wasteful and a few intermittent connecting welds will prove both more efficient and

economical.

3. If the plates being joined are to provide a pressure barrier—the steam pressure on the

two sides are different—then the welds must be constructed to produce isolation from

side to side.

In selecting the weld, it is necessary for the designer to address certain considerations

concerning the most appropriate form as follows.

• Are fillet welds acceptable and what size fillet is required? Will a single fillet be

sufficient, and is there access for a double fillet? If not, will a partial or intermittent

weld on one side be sufficient and acceptable?

• Are the stresses that will be developed at the weld shear, tensile, or compressive? The

stresses will dictate the type of weld and to some extent the sequence of welding.

• Are the stresses cyclic and is their magnitude sufficient to make fatigue a

consideration at the local operating temperatures? Is there a possibility of inducing

low-cycle fatigue into any of the joints?

• If the joint falls into the category of being subjected to a cyclic loading, which could

induce high stress values, what contour weld is required, and what surface finish must

be obtained on the welded surface and transition regions?

• Are fully penetrated welds required?

• Will partial penetration groove welds or fillet welds suffice as long as the full strength

of the plate is developed in shear or tension?

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The specification prepared by the design engineer for any fabrication will define the

applicable welding codes, their requirements, and how they will be met. The actual fabricator, or

manufacturing engineer, is primarily concerned with and is responsible for meeting these various

requirements of the designer.

The selection of joint geometry is the responsibility of the designer, who will select the

joint form, plate thickness, and the welding techniques and materials to be used. The designer

will evaluate and should be aware of the loads to which the fabrication will be subjected, and

also if these loads are direct or cyclic. Therefore, the fabrication specification will typically

include the following information.

• Any code requirements to be met. This will identify both the code and any special

inspection requirements by regulatory bodies.

• The joint geometry. The five basic joint geometries are shown in Figure 5–26.

(Blodgett 1956)(AWS Standard Welding Symbols) The engineer will normally define

the joint requirements and any weld preparation required. The designer must

determine if a fully penetrated weld is required, if a butt weld will be sufficient, and

whether the welds should be continuous or intermittent.

Fig 5–26 The Five Basic Weld Joints

• The filler material and weld processes to be used. This could include stick size and

the sequence of laying in the material.

• The fabrication sequence. The sequence is established normally after detailed

discussion with the manufacturing or production department. This is necessary to

ensure access and to ensure those welds requiring additional strength can be prepared

for complete penetration and if necessary can be back gouged for greater integrity.

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• Any requirement for preheat. Preheat which is not normally required in the

material used for turbine hood structures.

• Any stress relief requirements.

There are three types of weld—the fillet, the groove, and the plug weld. For turbine hood

fabrication only the first two need be considered. Typical joint configurations are shown

schematically in Figure 5–27.

Fig. 5–27 The Cross Section of Basic Weld Joints Used in the Fabrication of a Low-

Pressure Hood

Structural welding faults

In all structural welding, the opportunity exists for various forms of defect or fault to

occur. This is discussed further in “Whats Ailing That Weld?” in the Welding Journal, August

1997.

The result of these faults can be insignificant but can also be a source of mechanical

weakness causing weld distortion or high stress concentration in the various portions of the

structure involved. The most common faults are various forms of inclusion, cracks in the weld,

and undercutting:

Slag inclusions. Slag can be deposited within the weld as shown in Figure 5–28. These

inclusions can act as stress concentration centers. Such inclusions are captured in the base weld

deposit material during the deposition process. This is particularly common if the layers of weld

are deposited and the slag from one pass is not removed adequately before the next pass is

deposited.

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Fig 5–28 Slag Inclusion Both Single and Cluster

Oxide inclusions. Oxide inclusions occur when the surfaces to be joined are not cleaned

of oxide before the welding process begins. The oxide scale from metallic surfaces will, if they

have a sufficiently high melting temperatures, drop into the weld pool and become trapped.

These inclusions like those caused by slag can introduce high stress concentration.

Porosity. Porosity is the result of gas pockets becoming trapped in the weld bead. These

beads can exist as small single voids or in clusters. In either case, these voids represent

discontinuities that can introduce high stress concentration and have the capability to generate

cracks. The forms of these inclusions are shown as Figure 5–29.

Fig. 5–29 Porosity Single and Cluster

GTAW-tungsten inclusions. The use of gas tungsten arc welding can, if procedures are

not followed in detail, result in tungsten particles from the electrode being trapped in the weld.

Crater cracking. These are cracks that occur in the heat-affected zone (HAZ) and are

normally a consequence of not breaking the arc correctly. A method of preventing or minimizing

their occurrence is to reverse the arc back into weld bead while reducing the welding current. A

schematic of such a crack is shown in Figure 5–30.

Fig. 5–30 Crater Cracking

Hot and cold cracking. Hot cracking is normally caused by the inclusion of excess

amounts of sulfur or phosphorus in the steel or incorrect methods of breaking the arc. This is a

form of crack that occurs while the components being joined are at high-temperature caused by

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the welding process. These cracks tend to be transverse to the weld and are the result of excess

local shrinkage or excessive cooling rates at completion of the weld.

Cold cracking occurs after solidification of the weld material is complete and can initiate

days after the weld is complete. This is normally a result of hydrogen embrittlement. Cleaning of

the joint to remove moisture or any other liquid deposit before welding can help minimize the

occurrence of this condition.

Centerline cracking. A centerline crack is shown in Figure 5–31. This is a cold crack

that runs down the center of a single concave weld. This type of crack can initiate because the

weld bead is too small for the thickness of the plate, high joint restraint, the extension of a crater

crack, or poor fit up of the joint before welding proceeds.

Fig. 5–31 Centerline Cracking

Undercutting. Undercutting is shown in Figure 5–32. This occurs when the base metal

adjacent to the weld bead is cut or weakened at the intersection of the bead and base metal. The

result of this material removal is to reduce the load carrying capacity of the base metal and/or

increase stress concentration in that region.

Fig. 5–32 Undercutting

Incomplete fusion. This is a condition that occurs when the weld bead does not make

complete fusion to the base metal or even between passes. This is shown in Figure 5–33. The

principle causes for this condition are excessive travel speed, insufficient welding current, or

using too large an electrode. This condition is a natural location for high stress concentration to

occur and initiate major cracking.

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Fig. 5–33 Incomplete Fusion

Welding Consumables

As important as the selection and preparation of the weld geometry are the consumables

that will be used in the fusion process. They have a major impact on the integrity of the weld.

These consumables include the electrodes or filler materials, the fluxes and gasses and even the

rate and sequence in which these elements are applied. In the manufacture of the electrodes, their

diameter and chemical composition must be controlled within close tolerances to ensure they

meet with the requirements set by their own engineering.

The storage and treatment of the consumables is important as some consumables will

deteriorate if not maintained in a suitable condition. For semi-automatic and automatic welding,

the filler wires and fluxes must be controlled and stored with exactly the same level of care.

High-Pressure Shell Manufacture

The majority of steam turbine high- and intermediate-pressure shells are produced by

casting. The majority of these are from a high-quality alloy steel. The casting process is complex,

and because of the elaborate form of the casting, and the duty to which it will be exposed, there

is a need to minimize casting faults which can occur if proven procedures are not followed in

detail. The majority of the faults that occur must be identified in terms of size and location, and

must then normally be excavated and upgraded. Such upgrading is an expensive process and can

add considerably to the cost of producing a quality casting. However, this is an essential step to

providing a safe and reliable shell.

The initial step in casting manufacture of a high-pressure turbine shell is the production

of a suitable pattern from which the shells can be cast. Since the form of the final shells can be

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no better than the pattern from which they are produced, any effort and thought expended in the

production of the pattern, its design and manufacture can be easily justified. (Clymer 1968)

Patterns for large castings are often split vertically along their axial length to facilitate both their

manufacture and handling.

Some manufacturers produce a number of component pattern parts, from which a variety

of casting forms can be constructed for a variety of section and cycle arrangements. Also these

casings can, when necessary, provide any required internal configuration of the steam path. The

pattern must be made to help ensure the castability of the shell. This means arranging for steady

tapers in vertical walls, ensuring to the greatest extent possible a progressive solidification,

avoiding abrupt changes of section when this can be achieved, avoiding regions where stress

concentrations will occur, and using fluid form shapes rather than cubic.

Obviously, to achieve these requirements, a certain degree of compromise may be

necessary between those of the founder and those of the design engineer. Over the years of

experience gained in the production of such castings, a general agreement and knowledge of the

requirements of both parties has evolved, and an acceptable product can now be specified and

produced. It is usually possible to achieve a suitable balance and minimize potential fault

regions. Figure 5–34 shows the building of a turbine-casing pattern.

Fig. 5–34 Building the Pattern for a Casting of a High-Pressure Casing

The casting melt is prepared to the chemical specification of the designer, and the metal

is prepared by either the induction or electric furnace processes. When the melt is complete and

pouring begins, an analysis of composition is checked by making a ladle analysis from the initial

pour. It is also possible that by forced cooling an analysis of the ladle sample will be taken from

the furnace and completed before the pour begins to ensure the chemical composition is correct.

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The chemical composition must comply with the engineering specification within close

tolerances. This composition must be checked, approved, and recorded. A moderate variation of

carbon content beyond the engineering specified upper limit is considered acceptable by some

manufacturers, particularly when thicker sections are involved. Such excess carbon can help to

prevent carbon segregation in the casting and also help ensure a more homogeneous section and

achieve the specified mechanical properties throughout its thickness. However, such a decision

and the potential consequences of excess carbon must be carefully considered before acceptance.

If internal chills are used, their number, type, composition, and location in the casting

must be selected to ensure the required solidification characteristics are achieved.

When the casting has been poured, solidified, and cooled to about 750°F, it is shaken out

of the mold. At this point, it is normal to blast clean and anneal. The annealing temperature is

dependent upon the chemical composition of the casting. Annealing temperatures range in

specifications from about 1250 to 1750°F. The heat soak period usually depends upon the casting

thickness, but is very rarely less than 10 hours. During this annealing period, a deviation of

furnace temperature of more than about 75°F is unacceptable. At completion of the heat soak, it

is normal to let the casting cool in the furnace to a temperature in the range 400 to 575°F.

At this point the casting can be removed from the furnace and the heads burnt off. Care is

taken to maintain the casting temperature. In burning off the heads, the casting main body is not

burnt. Remaining portions of the heads are removed by grinding and chipping.

If the casting is to be subjected to further heat treatment, the test blocks that are produced

integral with the casting are left attached. If not, the coupons are removed at this time.

At this stage in the manufacturing cycle, the casting is allowed to cool. All loose scale,

metal flakes, and core irons are removed. Blasting and grinding prepares the casting for NDE. It

is normal at this juncture for manufacturers to check wall thickness using ultrasonic methods

and/or templates. Rough machining is then completed. Such rough machining involves removing

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material from the horizontal joints and skim cutting the inner surfaces. This rough machining

leaves sufficient stock for the final close tolerance finishing. Rough machining is usually carried

out on individual halves of the shell, and no great effort is taken at this stage to match the two

halves.

Figure 5–35 shows the rough machining of the lower half of a nuclear turbine high-

pressure shell on a numerically controlled boring machine.

Fig. 5–35 Rough Machining the Lower Half of a Cast Shell for a Nuclear Unit

Before any detailed NDE is carried out, the casting is visually examined for gross casting

errors and faults. If this visual inspection indicates the casting is acceptable, it is then given a

magnetic particle inspection followed by ultrasonic and/or radiographic examination. During this

examination, special attention is given to areas designated for planned fabrication welds since

faults in these areas are greater cause for concern and more difficult to tolerate.

Faults detected during NDE are reviewed and evaluated. A disposition is then made

regarding acceptability and any corrective action required. This is necessary because faults must

be considered in conjunction with the temperature and predicted stresses in the region where they

are present. Also they must be evaluated in terms of the possible consequences of thermal

cycling. General guides for various faults are given in Tables 5–8 and 5–9.

Table 5–8 Faults Normally Unacceptable Without Rectification

Table 5–9 Faults Normally Acceptable Without Rectification

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Certain minor faults removed by grinding can be accepted without weld repair if wall

penetration has not been too great. If the fault depth is not excessive, the resulting crater should

be blended. It is normal and good practice, and in some jurisdictions a legal requirement, to

record the extent of excavations with photographs or sketches. Such a record photograph is

shown in Figure 5–36.

Fig. 5–36 An Upper Half Shell for the High-Pressure Section of a Fossil Unit, Showing

the Excavations Required to Remove Casting Faults

The casting is upgraded, and if required, the normalizing and tempering cycles are

completed. These processes are undertaken either with the test coupons still attached to the

casting or, if previously removed, included in the furnace.

Casting upgrading by weld repair is an acceptable procedure for returning the shell to an

as-designed condition. This total procedure includes the preparation of repair procedures, heat

treatment, examination, and testing after excavation to ensure all traces of the defect have been

removed. Final examination and NDE at completion of the weld repair are also required.

The introduction of X-ray equipment for the examination of castings has increased the

capabilities of manufacturers for detecting and evaluating faults. The facility, Figure 5–37 shows

a 10 million volt linear accelerator used by one turbine casing manufacturer to detect casting

faults. (Kent 1964) This unit has the capability of penetrating an 18-inch thickness. Taking such

X-rays requires careful judgment on the part of the operator to ensure an acceptable and

interpretable quality picture is produced. Underexposure or overexposure can easily lead to a

fault going undetected. Because the exposure is a function of the applied voltage—usually fixed

for any piece of equipment—and exposure time, careful control is required.

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Fig. 5–37 A 10 Million Volt Linear Accelerator Used to X-ray Turbine Casing Shells

The final or precision machining of the shells is undertaken after all upgrading and heat

treatment is complete. Such final machining is normally completed with the half shells connected

(bolted) together as seen in Figure 5–38. This requires the horizontal joints be finished first and a

checks indication that joints are tight. This requires the use of an engineer’s blue mark.

Simultaneous boring of the upper and lower halves ensures any machining mismatch is

eliminated and helps ensure that casing sag is closer to the operating condition.

Fig. 5–38 The Half Shells Bolted Together for Final Machining

Therefore, it is essential when undertaking this machining that the castings are supported

on the machine tool in a manner similar to the means of support provided when the unit is

mounted in the field. While missing the temperature effect, supporting the casing by any other

manner could introduce an unrepeatable sag that could not be duplicated in the field. This would

lead to a degree of non-concentricity of the various bores and misalignment in operation. At

completion of final machining, the shell halves are hydrostatically tested. In Figure 5–39 is

shown a large half casting for a single cylinder unit after final machining.

Fig. 5–39 A Shell Casting After Final Machining

Shell Casting Faults

Due to the high melting temperature of casing steel, and the fact that the material will not

cool evenly throughout the shell form, alloy steel castings have a high probability of developing

certain forms of defect. (Rogers and Brewer 1964–65) The shells are the largest castings used in

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the unit and they are subject to the highest levels of stress in operation so they are also those

elements demanding the greatest amount of attention in upgrading.

The nature of defects most likely to occur allows them to be placed into five major

categories. These are:

1. shrinkage cavities

2. cold cracks

3. hot tears

4. porosity

5. scabs

We will consider these five types of defects separately.

Shrinkage cavities

As their name implies, these defects are voids formed in the steel when the casting cools

through the liquid, solidification, and solid phases. Castings cool from the outside with a solid

phase achieved first in the outer skin. If there is insufficient feeder head available to the internal

portion of the casting, cavities will form as further solidification occurs at the outer surfaces. It is

important during the design of the pattern that headers of sufficient volume are provided and that

these are located so they can supply liquid metal to the casting as it cools and have a sufficient

reservoir after the pour.

Since shrinkage cavities are the result of slower internal cooling, it is only rare that they

occur at the surface of a casting. Figure 5–40 shows an excavated shrinkage cavity exposed by

machining.

Fig. 5–40 Shrinkage Cavities Exposed by Machining

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Cold cracks

This form of defect occurs in the cooling casting at lower temperatures and is the result of

internal stresses set up in the shell by contraction. As the casting cools, solidification will occur

in different portions of the total form, and there will still be metal in the liquid phase between

such solidified regions. Then, as these final liquid phases cool, there is solidification. If this

solidifying material is unable to adjust for the different levels and direction of cooling and

shrinking, then internal residual stresses are produced. These stresses tend to increase in

magnitude with further reduction of temperature and shrinkage.

As the temperature of the shell casting falls, the stresses eventually exceed the ultimate

tensile stress limit of the material at its local temperature and a crack is initiated. A cold crack at

the surface of a shell is shown in Figure 5–41.

Fig. 5–41 Cold Cracks at a Ground Surface

Hot tears

This form of tear is similar in many respects to cold cracks, except these form due to

excessive stress concentration at temperatures just below the solidification temperature. In

comparison to cold cracks, these cracks tend to be generally irregular and jagged in appearance.

Cold cracks form principally at the surface of a casting. However, hot tears may be either

internal or at the surface.

Due to their irregular nature and the fact they may have many branches, these tears are

often difficult to locate, particularly when they exist just below the surface. Figure 5–42 shows a

hot tear in a shell casting. Such internal tears are best located by means of radiographic

examination.

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Fig. 5–42 Hot Tears in a Shell Casting

A phenomenon experienced on some castings is known as filamentary shrinkage cracks.

These occur at about the center wall thickness and consist of a concentration of many fine

cracks. These cracks, if not found and removed, will work outward to the surface after periods of

operation at high stress levels.

Porosity

This type of defect is the result of hot molten metal contacting damp sand and generating

steam. This steam is immediately oxidized causing the formation of hydrogen gas. If this

hydrogen gas is unable to escape through suitable vents, it will become trapped in the casting,

forming voids or gas holes that may be pin-hole size or the more typical blow holes.

To prevent this type of fault, it is important to dry the mold before pouring, and to ensure

the steel has a low hydrogen content. It is also necessary to provide adequate venting to allow the

gas that is generated during pouring to escape.

Scabs

A mixture of sand and metal exists at the surface of many castings. This mixture is

extremely hard and difficult to remove. The scab mixture is caused by a combination of hot

shrinkage of the sand and spalling—a flaking of the surface sand. Additives can be included in

the sand to improve surface binding intended to minimize the possibility of this effect.

The Upgrading of Castings

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The various faults and the regions in which they can occur in a shell casting make it

necessary to develop procedures by which castings can be repair welded, or upgraded, after the

faults have been located and excavated. Because the heating cycles, through which the casting is

cycled in operation, are similar to and in some cases even more severe than its casting process, it

is necessary to develop stringent process instructions and then control the methods employed to

cover these upgrading methods. These repairs should be made by qualified welders.

One of the primary concerns with reheating a casting in preparation to undertaking

repairs is that the casting surface will form a scale. A procedure used by some manufacturers is

to paint the casting on any surface that could be affected by scaling. A paint that will maintain

surface coverage at high-temperature and minimize this effect is used.

Faults are located by NDE and the casting marked to indicate their location. These faults

are then excavated. Such excavating requires the area be explored and all the affected material

removed. Figure 5–43 shows a casting for the top half outer shell of an intermediate pressure

cylinder. Shown are the radiographic grids and chalk marks around discovered faults and the

extent of the excavations required to remove them. Surface faults in one area have been ground

away; these may or may not require repair depending upon the depth of excavation necessary to

remove them and the thickness and form of the remaining material.

Fig. 5–43 Half of a High-Pressure Nuclear Shell Casting Showing the Marked Results of

NDE and Indicating Where Excavation and Upgrading is Required.

To permit welding repairs, the casting must normally be preheated, and this preheat must

be maintained throughout the repair procedure. Typical alloy steel materials for castings require

a preheat in the range 475°F to 850°F. Such preheat can be localized or total. For high alloy

steels, total preheat is preferred to minimize the possibility of crack formation. When it is not

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possible or practical to uniformly preheat the entire casing, it is necessary to surround the repair

area and heat for a distance of 12 to 16 inches on all sides of the repair. The rate of heating the

affected area should be so controlled that no part of the affected zone shall be hotter than any

other by more than about 100°F.

During the deposit of weld filler material, the preheat flame should be kept in constant

motion over the repair zone. In no case should the temperature of the casting be allowed to

exceed the nominated preheat temperature range.

During the welding process any cracks, slag inclusions, undercutting, porous crater,

granular flux, or poor fusion regions that appear on the surface of a pass must be removed before

depositing the next pass. It is often justified to make a magnetic particle examination at the

completion of some weld passes to ensure no cracks have developed. At completion of weld

deposition, the area should be allowed to cool at a rate not exceeding 100°F/hr.

At completion of the total weld repair, the rebuilt area is dressed. The casting is then heat

treated for stress relief. It is necessary to record the thermal history of any heat treatment that is

undertaken. A record of such a repair heating operation is shown in Figure 5–44. If this stress

relief operation is not complete and any internal residual stresses are not removed, there will be a

tendency for the casting to warp and distort upon machining. Such residual stresses may later

aggravate operational stresses and affect unit alignment.

Fig. 5–44 Data Produced From a Thermal Histogram for the Heat Treatment of a Turbine

Shell Half

When repairs are made to a chrome moly-vanadium casting, it is often expedient to use

chrome moly repair rod. This is necessary because of the difficulty of producing crack free welds

with electrodes containing vanadium. It is normal to check the mechanical properties of the weld

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rod used for repairs, because these areas are often the cause of operating problems if mechanical

properties are not maintained.

Shell Manufacturing Tolerances

Since the casing contains high-pressure, high-temperature steam, it is necessary to ensure

all steam joints provide a good seal to prevent leakage. Leakage is efficiency degrading and

potentially dangerous. It is capable of causing serious injury or death to operations personnel.

An effective steam-tight joint is the product of two factors—a pair of flat surfaces and an

adequate bolting or clamping system that can pull the surfaces into hard close contact without

exceeding allowable stresses in the bolt or clamp. Modern manufacturing techniques allow the

horizontal shell joint to be produced by machine tools that require little or no handwork. That is,

the surface finish and flatness are normally acceptable as produced.

The inner shells for high- and intermediate-pressure cylinders should be closed to within

0.0015 in. without studs and bolts assembled. In addition, engineering blue marks will indicate

an 80% marking in high-pressure or temperature regions with a continuous band over the entire

periphery of not less than 1.0 in. wide at any point, with steam conditions above either 2200 psi

or 750°F. Below these conditions, 40% markings should be indicated with a continuous band of

0.75 in. minimum width at any point.

The requirements for outer casings are less stringent, requiring a maximum 0.006 in. gap

without bolting and 0.0015 in. with every third bolt nipped. Engineering blue marks will be 40%

overall and 80% in the region of the glands. The horizontal joint face should have a maximum

surface finish of 125 micro-inches ( _m-in.) on both inner and outer shells.

If the shells have grooves in them that will be used for locating diaphragms and are

therefore required to produce a steam seal face, these axial faces should have a finish in the range

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64 to 125_m-in. If stationary blades are located directly into the casing, the finish will be

dependent upon the requirements of location and possible bending stresses that will be induced.

The Joining Of Casings Parts

There are various basic designs of casing used to contain the high-pressure high-

temperature steam. These alternate designs are selected to minimize stress in the various

components while at the same time making the casing as thermally flexible, in terms of its ability

to accept and reject heat, as possible. Such flexibility is achieved using various forms of casing.

However, these designs consist of three main types, and the difference between them is

the method employed to provide access for the rotor and then join together the various casing

halves with sufficient security they can operate without introducing excessive leakage at the

positions where the joints are made. The three principle methods of access and joining follow.

The bolted horizontal joint

The majority of high duty casings are joined at their horizontal centerline by the use of

threaded components connecting the two halves together at a flanged joint. This horizontal joint

represents a discontinuity in the thickness of the casing walls and does not have the same ability

to accept and reject heat during transient conditions. However, it does add considerable rigidity

to the casing and also helps maintain alignment because of the stiffness of this mass.

The two casing halves are joined by bolting or, more correctly in the majority of designs,

by the use of studs. These studs are screwed into the lower half and tightened by means of nuts

connecting and holding the top half casing in intimate contact with the lower half. The design

process analyzes the total load that is developed inside the casing by the high-pressure steam,

and then selects a bolting pattern sufficient to hold the casing halves in tight contact.

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This clamping must be achieved without the individual studs exceeding stress levels that

would cause failure or make their useful life unacceptable in terms of the number of hours of

operation and the number of times these elements can be removed and reused.

Shown as Figure 5–45 is the bolting pattern for the half portion of a high-pressure, high-

temperature casing design employing inner and outer shells. In this layout, each bolt hole is

numbered and the bolt size required at each location defined. It is a normal practice with some

operators to maintain a log of the duty of the connecting stud at each location. They log the

number of hours of operation and the number of times the stud has been removed from service

and re-tightened.

Fig. 5–45 The Bolting Pattern for One Symmetrical Half of the Inner and Outer Half

Casing of a Fossil Unit

From this information, it is possible to predict when studs should be changed and then to

have replacement elements available for installation at a suitable outage.

An important consideration with this type of design is the rate at which heat can be conducted

through the horizontal flange to modify the temperature of the studs so they can achieve a

compatible temperature and not be over-stressed by the heating or cooling of the casing. Studs

must be able to accommodate the thermal expansion and contraction loads induced by steam

temperature changes.

Similarly, it is important that in specifying the bolting pattern there is sufficient material

remaining in the flange, in terms of the diameter and distance between hole centers, that stress

levels in the casing flange do not become excessive.

Many older designs employed grooves produced on the face of the horizontal joint, which

is designed to pass high-temperature steam between the bolt holes so the studs are able to

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achieve operating temperature at a faster rate in response to major temperature changes in the

steam.

The shrink ring joint

A second design joining the casing halves is that employing shrink rings. With this

design there is a combination of studs and shrink rings used to make the joint. While the studs do

make some contribution to the total joint strength, their major function is to hold the two halves

in their correct position while the shrink rings are heated and assembled over the casing halves.

Shown as Figure 5–46 is the section through a high-pressure unit in which the inner

casing is held together by a series of nine shrink rings placed along the axial length of the casing

halves. Eight of these rings are on the downstream side of the nozzle box and one is placed on

the upstream side and above the balance piston.

Fig. 5–46 An Inner Casing With Shrink Rings

Shown in Figure 5–47 is the detail of a shrink ring dimensional requirements. Here the

casing surface is machined to a diameter of Dcc in the cold condition. In this cold condition the

shrink ring has an inner diameter of Drc. Therefore, when the ring is heated and assembled to the

casing it will upon cooling produce an interference fit of _ic=Drc-Dcc. The amount of shrink

interference fit is selected by the designer so the ring will clamp the halves but not be strained to

the extent there will be any plastic deformation of the ring.

Fig. 5–47 Details of the Shrink Fit

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When steam is admitted to the unit, both the rings and casing will heat and expand. The

rings will be surrounded and attain the temperature of the steam discharge from the high-pressure

last stage blades. There will also be some small amount of heat transferred by conduction

through the casing. The internal steam temperature is higher than that of the steam surrounding

the rings.

The casing will attain an equilibrium temperature between the steam space gap and inner

wall temperature. At these higher conditions the casing will expand to Dch and the ring to Drh.

There will under these conditions be an interference fit of _ih=Dch-Drh. Again, in the hot

condition, the shrink fit will achieve an interference sufficient to maintain contact between the

two halves but not so tight as to cause plastic deformation or extensive creep of the shrink rings.

The process of heating the shrink rings enough to expand them is shown in Figure 5–48,

where a gas ring is located around the ring, heating it at a constant rate to a constant temperature

sufficient to allow assembly.

Fig. 5–48 Heating the Shrink Ring Prior to Assembly

The end-loaded rotor design (barrel construction)

The end-loaded design is shown in Figure 5–49. Here, the inner casing is a bolted

construction with the rotor and stationary blades assembled in the normal manner and bolted at

some horizontal joint. However, this inner assembly is end loaded into an outer casing that has a

close to cylindrical form. The outer cylinder is closed by the bolted attachment of an end cap,

which produces a steam tight joint with its internal pressure defined by the exhaust pressure from

the last stage blade of the section.

One significant difference with this design occurs when it is necessary to open this

section. It is necessary to disassemble any couplings to adjacent sections, then either tilt the outer

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cylinder and remove the inner assembly or withdraw the inner casing through the open end. Such

disassembly will be required each time it is necessary to inspect or undertake work on the

internal cylinder or steam path. With correct training, this process can be accomplished with the

same time span of a normal maintenance outage and should not impact adversely on schedules.

Threaded Components

A major application of these components is in the connection of the joining of casing

halves. These threaded connectors used for joining the casings are intended to provide enough

contact to form a steam tight joint that is able to maintain the seal for extended periods. Such

components must also be able to be disassembled and then rejoined and able to maintain the

efficiency of the joint upon reassembly. The fasteners must be able to withstand high levels of

tensile load without failure.

To define the requirement of threaded components it is necessary to define characteristics

that establish their form and that can be used to gauge their acceptability. (Chapman 1958) These

are shown in Figure 5–50.

Fig. 5–50 Characteristics Defining the Screw Thread

Full or major diameter. For a straight thread, the major diameter is the diameter of a

coaxial cylinder that would bound the crests of an external thread or the root of an internal

thread. For a tapered thread, the major diameter is the diameter of the major core at that position.

The core or minor diameter. On a straight thread, the minor diameter is the diameter of a

coaxial cylinder that would bound the roots of an external thread and the crests of an internal

thread. On a tapered thread, the core diameter at a given position is the diameter of the minor

core at that position.

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Effective pitch diameter. On a straight thread, the pitch diameter is the diameter of a

coaxial cylinder the surface of which passes through thread profiles at points that would cut them

at a width where they are one-half pitch thick. On a perfect thread, this effective diameter is the

mean of the full and core diameters.

Pitch. The distance measured parallel to the thread axis between corresponding points on

consecutive contours.

Thread angle. This is the angle subtended by thread flanks measured in an axial direction.

Crest radius. This is the radius joining adjacent flanks at the thread crest.

Root radius. This is the radius joining adjacent flanks at the thread root.

To achieve interchangeability of any threaded component, these dimensions must be the

same on both. These seven dimensions, together with flank flatness, influence the strength and

load carrying capability of the thread. Of these seven dimensions the most critical are the pitch,

thread angle, and the pitch diameter.

Forms of pitch errors

In the manufacture of a screw thread, the most critical consideration is maintaining the

correct pitch. Any excessive errors can cause uneven pressure between the flanks and affect

sharing of load between adjacent profiles. This can lead to local yielding and eventually failure at

stresses and loads well below the load carrying capability of a conforming element. There are

several sources of pitch error, the most common being introduced during the metal cutting

process due to tool or cutter wear or material stock being incorrectly clamped.

If the component is processed after the initial thread forming operation, these processes

may also introduce errors. Any process requiring the application of heat can cause errors and

such errors may be immediately apparent. However, they may also occur after a period of

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relaxation has occurred in the threaded component. Further manufacturing processes for

finishing such as lapping, plating, and grinding may also introduce errors.

The most common manufacturing processes used to produce threads can introduce three

common forms of pitch error.

Progressive errors. This type is a uniform error in pitch resulting in a progressive

lengthening or shortening from the nominal. If a large number of pitches are in engagement, this

will eventually cause thread binding when assembled to a component with conforming pitch.

Periodic errors. Periodic errors are those that occur at intervals. They may produce a

cumulative error or they may be self correcting and subsequent threads be at the correct axial

location relative to the first pitch.

Erratic errors. These errors can cause either a shortening or lengthening from the

nominal. They are not periodic and are unpredictable.

Bolt stresses

The stresses that occur on threaded portions of a fastening are of two types, tensile and

shear, it is also possible for a bolt or stud to be subjected to a torsional load. However, this is

generally not the intended purpose of such a component, and if it occurs, represents either poor

design practice or wear causing the component to be loaded in a manner for which it was not

designed. In the turbine shells, the bolt are designed for and are loaded in tension.

To estimate the tensile or shear stresses that occur in the threaded component, it is

necessary to be able to determine the effective area of the threaded load-carrying portion. This

can be found from the expression:

Ab = 0.7854 D - 0.03937

n (5.4)

where

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Ab = effective area of the threaded portion

D = full or nominal diameter of threaded portion in millimeters

(mm)

n = number of threads per mm

The nominal or mean stress is a good indication of the levels within the thread elements.

It does not, however, represent the actual stresses that occur. There are two factors that act to

increase the actual stress levels above the nominal.

1. The complex form of the thread and the fact that load transmittal is through thread

teeth, which can deflect under the action of the applied axial loads.

2. The pitches of the threads are not exact. Therefore, it is not possible to ensure the

individual portions of the thread engage together to the extent the load will be shared

evenly.

These two effects can be amplified when threaded components that have operated at

high-temperature are reused, particularly if there has been any scaling during the previous

operating period. It is good practice to ensure that when such components are reused, the stud or

bolt is returned to service employing the same nut as used in the first application.

If a new stud, bolt, or nut is used in conjunction with a used component, there will

normally be some readjustment as the assembly is tightened and returned to service.

The types of failure in threaded components

There are three major types of failure that occur in the threaded components.

In the male thread at the nut face. Figure 5–51 shows a tightened nut and bolt and their

stress or load transfer lines. It is clear the nut reverses these lines and converts the tension in the

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bolt shank to a pressure between the nut and face of the component being compressed. It is also

clear from this, there is a potential for large stress concentration in the threads and shank under

the nut head. In fact the stress concentration in this area is about 4.0.

Fig. 5–51 The Stress Lines in a Tightened Nut and Thread

Rupture at the nut face. This is the most common form of failure, and accounts for

about 65% of the total failure in threaded components. (Peterson 1973)

The stress concentration factor at the interface can be reduced to about 3.0 by designing

the nut as shown in Figure 5–52. In this arrangement, the peak stress is reduced by virtue of the

lip being stressed in the same direction as the male portion. The stress lines of such a

combination is shown in Figure 5–53.

Fig. 5–52 A Modified Nut Design to Reduce Stress Concentration

Fig. 5–53 The Stress Lines in the Modified Nut

The use of a nut produced from a material of a lower modulus of elasticity is helpful in

reducing the peak stress in the male threads. The less rigid material of the nut elastically deforms

bringing more load bearing nut threads into contact with the male threads. Thicker nuts may also

help in reducing peak stress by providing more threads to share the total load.

Many high-temperature applications employ a tapered bolt in an effort to achieve a better

load distribution over the bolt threads. A suitable taper rate is 0.003 in. to 0.008 in. per inch of

threaded bolt length.

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A major factor contributing to failure under the bolt head is fatigue. Many threaded

components are subject to cyclic loading. Such cyclic loading causes fatigue failure that can be

aggravated if the nut contact face is not flat. Tests have indicated if the nut contact face is shaped

as in Figure 5–54, then life is reduced. If _ is 0.5 degrees, fatigue life is reduced by 20 to 50%

and a 1.0 degree taper reduces life by 60 to 80%. For critical applications, nut flatness is

therefore of considerable importance, and variance from the flat in excess of 0.10 degrees is not

acceptable.

Fig. 5–54 Nut Head with an Angled Underface

A failure surface from a stud is shown as Figure 5–55, which is a 2.5 in. diameter bolt

from an inner casing. This stud has been subjected to an operating temperature of 800°F. An

examination of the surface indicated the fracture initiated due to creep at the root of the thread,

possibly from a small crack initiated by impact load.

Fig. 5–55 Failure Surface Under a Nut Head

Rupture at the last thread on a shank. This is the second most common form of failure

and accounts for about 20% of the total. Figure 5–56 shows the failure surface at a vanishing

(last) thread on a stud.

Fig. 5–56 Failure at the Last Thread

The reason for the high failure rate in this region is that the stress concentration factor

—3.30 to 3.40 for a normal thread—becomes 4.40 to 4.50 for a washout or vanish thread, with

the actual value depending upon details of the thread geometry.

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The effect of stress concentration under the end thread is shown in Figure 5–57. This

effect can be reduced by designing the shank to have a reduced diameter about equal to the

thread core diameter at the termination of the threaded portion. This effect is shown in Figure

5–57b.

Fig. 5–57 The Lines of Stress Concentration at the Threads

In general the fatigue endurance limit can be increased by the use of finer threads. For

example, a 0.75 in. diameter shank with 30 threads per inch will have a 25% increase in

endurance limit over one of the same diameter with 15 threads per inch.

Under the bolt head. This is the last of the common forms of failure. Failures in this

region account for about 15% of the total. Consider the diagrammatic bolt head shown as Figure

5–58. There is heavy stress concentration at the contact point. This concentration factor may be

of the order 5.5 to 6.5 the actual value depending upon local geometry.

Fig. 5–58 Showing the Stress Lines and Stress Concentration at the Contact Points

The effect of stress concentration can be reduced somewhat by the use of elliptical or

parabolic fillets under the head. Increasing the head thickness h also helps reduce the bending

stress and the concentration at the interface.

Stress relaxation

The high stress levels in many threaded components give rise to some relaxation during

operation. This is particularly so when the part is operating at an elevated temperature. This

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effect will modify the overall performance of the bolt and its ability to clamp. As a consequence

of this, many parts may require a periodic retightening.

Stress relaxation is considered to be a function of both time and temperature. Tests

indicate an initial loss of load of from 2 to 11% is normally experienced immediately upon

completion of the torquing. The average loss is about 5% of the maximum registered bolt

tension. (Fisher 1974)

This drop in tension is believed to be due to elastic recovery that takes place when the

torque-wrench is removed. Creep and yielding at the thread root due to high stresses may also

contribute to minor relaxation and possibly redistribution of the total load.

Relaxation tests have shown that after the initial relaxation, a further 4% loss in the male

portion occurs in a period of days when compared to measurements taken one minute after

torquing. About 90% of this relaxation occurred during the first day. During the remaining days,

the rate of change of bolt load decreased in an exponential manner.

Temperature has an adverse effect upon and contributes to relaxation. The following

residual stresses were noted in CrMoV steel bolts after 10,000 hours operation at elevated

temperatures. The initial stress is 36,000 psi.

Temperature (F) 800 900 1000

Residual stress (psi) 23,350 16,250 5000

Preloading of bolts

When a bolt is pre-loaded it is constantly under a tensile stress. During operation, these

stresses may be increased. Therefore, if initial stresses are too high the bolt can fail or the

material can exceed its yield stress.

Many threaded components of the steam turbine are subject to compound loads imposed

initially by the tightening process and compounded by the effort required to hold the flanges

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together against an internal steam pressure. During operation, the bolt is also at a much higher

temperature. This temperature effect reduces the mechanical properties of the material, which

may also have imposed on it additional stresses due to heating and cooling of the turbine parts.

Consider the bolted joint shown as Figure 5–59. Assume this bolt is one of a series holding

together two components subject to internal pressure. Further, consider the effective part of the

casing flange influenced by the bolt is a portion designated as Ac.

Fig. 5–59 The Forces Induced in a Threaded Joint by Tightening and Internal Loading

Upon initial tightening, a force f, being a function of the ratio of the effective area of the

bolt Ab to the area of the flange associated with the bolt Ac, is applied to the two components.

This force f is also minimally influenced by the ratio of the moduli of elasticity of the different

materials between the bolt and flange. In Figure 5–59, the loads to the left of the centerline are

those due to bolt tightening and those to the right are those after steam loads are applied.

After initial tightening:

Extension of Bolt = f . L

Ab . Eb and, Compression of Flange =

f . LAc . Ec (5.5)

where

L = effective length of bolt or stud

Eb = modulus of elasticity of bolt material

Ec = modulus of elasticity of flange material

After the steam load is applied a new force is developed on the bolt of 'F due to internal

steam pressure P where P < f.

Therefore Extension of Bolt = F . L

Ab . Eb

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and (5.6)

Decrease in Compression of Flange = ( F - P ) . L

Ac . Ec

Therefore the increase in bolt extension εb due to P is:

Increase in Extension of Bolt =

( F - f) . L

Ab . Eb

and (5.7)

Decrease in Compression of Flange = ( F - P ) . L

Ac . Ec

Equating these gives:

F = f + P

1 + Ac . Ec

Ab . Eb . L (5.8)

For the major components of a turbine subject to stress, the ratio Ac/Ab is between 5 and

10. For example, for the casings of a high-pressure fossil-fired unit the ratio is about 6. If it is

assumed that Ec/Eb =1, then:

F = f + 0.14 P (5.9)

Using this value of F, and knowing the effective area of the bolt Ab the nominal stress

can be calculated.

Considering the simplest case, extension of the bolt εb can be found from:

εb = F . L

Ab . Eb

(5.10)

Therefore, the stiffness constant Kb of the bolt can be found from Kb = F/εb.

That is Kb =

Ab . Eb1 (5.11)

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If washers or other components are gripped between the bolt, then their stiffness must

also be considered. In this case the total stiffness Ks can be found from the relationship:

1Ks

= 1Kb

+ 1K1

+ 1K2

+ 1K3

. . . . 1Kn (5.12)

In equation 5.12, K1 … Kn are the stiffness constants of the other components being

compressed by the bold as it is tightened. If one of these components is a soft gasket, its stiffness

relative to the other members is usually so small that for practical purposes the others can be

neglected and only the gasket stiffness need be considered.

Reuse of high strength bolts and studs

Bolts and studs due to tightening and internal stresses during operation often have

stresses induced in them that exceed the elastic limit of the material. The repeated tightening of

high-strength bolts can, under these circumstances, be undesirable. The records of a test on

repeated tightening of such a bolt causing stresses beyond the elastic limit are shown in Figure

5–60. It is apparent the cumulative plastic deformation has caused a decrease in the bolt

deformation capacity after each succeeding tightening, which was beyond the elastic limit in

each case.

For this reason it is necessary to control the use and stress levels of all bolts and studs so

as to maintain their load carrying capability.

Fig. 5–60 The Effect of Successive Tightening on the Life of a Threaded Component

The tightening of large bolt and stud elements

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It is known that the correct application of tightening procedures can extend the useful life

of threaded components. For this reason, it is important to ensure the correct degree of tightening

is achieved as this will produce joints which are both serviceable and capable of being remade a

number of time using the same threaded components.

The most accurate manner of ensuring the correct tightening of the bolt is to measure its

extension during the tightening process. The extension required in the bolt or stud can be found

from:

Required Extension = Pre stress x Effective Length

Eb (5.13)

During some tightening processes, it is possible the extension cannot be measured. If

such a process is employed, the nut or bolt head advance can be calculated from a knowledge of

the thread geometry. The required length of arc is found from:

C = χ . π . D . e

P (5.14)

where

C = length of arc

D = outside diameter of nut

e = extension required

P = pitch of thread

_ = the tightening factor = _1 + _2

The factor _ is the initial tightening factor, characteristically equal to about 1.4. It

comprises two components _1 and _2.

χ1 = 1.2 =

Af + AbAf (5.15)

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where

_2 = 0.2 = bedding down factor

Ab = effective area of bolt or stud

Af = effective area of flange

Threaded component material

For the majority of bolting materials in the steam turbine horizontal joint positions, the

material used must be selected to achieve certain mechanical properties at high-temperature and

for continual operation at high stress levels. For this reason, bolting materials are produced to

stringent specifications and normally each turbine manufacturer will have developed materials

applicable to his unit and suitable to the type of application to which the components will be put.

An important consideration with high-temperature bolting in particular is that if bolts,

studs, or nuts are replaced, it is important the replacement elements use the same class of

material. A different material could have coefficients of expansion different from the original

and this is would cause loads to be shared unevenly when there is a temperature transient in the

unit, and some elements will bear an unacceptably large portion of the total load—even

sufficient to cause their failure.

The chemical composition of typical bolting materials is given in Table 5–10 and typical

mechanical properties in Table 5–11. These tables also indicate the temperature range over

which these materials are most often used.

Table 5–10 Chemical Composition of Typical Bolt and Stud Material

Table 5–11 Mechanical Properties of Typical Bolt and Stud Material

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These listed mechanical properties and chemical constituents represent the generic type

materials each manufacturer will develop specifications for their particular type of application

and will specify heat treatment to achieve the material properties required. These materials

should be free from cracks, surface flaws, or laminations. If there is any evidence of these

defects, the component manufacturer should reject this as unsuitable for the intended application.

The material specification must define the extent or degree of any defect determined by a defined

nondestructive examination.

Pipe Connection Points on the Casings

The steam entering both the high- and intermediate-pressure casings contains a

considerable amount of thermal energy with a high potential for leakage, and the joints are

exposed to considerable stresses. Such leakage can represent a considerable energy loss and a

degradation of cycle efficiency. It is also possible that certain of these joints will need to be

disassembled at maintenance outages. If there were no effort on the part of the inlet pipes to

move relative to the casings under conditions of start-up, shutdown, and during thermal

transients and no large thrust forces were developed, it would be sufficient to attach the inlet

pipes to the casing by any suitable and convenient method.

However, during warming, cooling, and condition transients, there is often a sufficient

change in the rates of expansion between the pipes and casing that there is relative growth

between them sufficient to cause a force in the pipes that can move the turbine and disturb the

alignment between stationary and rotating parts. Therefore, provision must be made to

accommodate this movement.

Some of this potential movement is accommodated by the methods used to locate and

hang the pipes, while other connections must be designed to allow relative movement at the

piping/casing interface sufficient to limit leakage and also to allow disassembly.

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There are several methods of conveniently introducing steam into or removing it from the

casing with each providing flexibility and minimizing steam leakage losses. There are four basic

systems that may be used individually or in combination to allow the pipes to connect to and if

necessary pass through the outer casing, then enter and connect to the inner casing or nozzle box

sufficient to allow the steam to enter the first stage stationary blade row. These methods are

discussed next.

Bolted connection

The simplest form is that in which a flanged and bolted connection is made. This type of

connection is used for lower rating units where the initial steam pressures and temperatures can

be classified as moderate. Such a connection could not be used for modern high steam condition

units. However, such a connection can still be used effectively for steam extraction points.

When such flanged connections are made, the flanges are arranged so the inlet pipes can

be disconnected to allow removal of the casing. This requires a second connection be made

outside the bounds of the casing to allow vertical withdrawal as seen in Figure 5–61.

Fig. 5–61 The Lifting Gap Required to Remove the Top Half Outer Casing

In such a system, it is necessary to ensure there is sufficient flexibility in the pipe system,

and the pipe hangers are maintained to retain piping flexibility.

Upper half weld

A common type of connection made to many higher rating units in the top half is the

welded form. This connection is used extensively in both high- and intermediate-pressure

sections. To allow connection, the casing is cast with stubs integral with the main shell casting.

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In this type of connection it will be necessary to make some other type of connection remote

from the welded joint connection on the casing to allow disassembly and vertical removal as

discussed for the previous type. Again piping flexibility is important.

Lower half weld

Similar to the welded connection discussed previously is the welded connection on the

lower half shell. Again, there is a need for a stub to be produced integral with the casing. Such

studs can be seen in Figure 5–35. These connections vary from the second type only in that it is

not necessary to connect second joints to allow withdrawal of the casing. To help ensure the

forces produced on the casing are minimized, and to simplify the piping layout, it is normal to

make these connections in the vertical direction.

In both of the welded-type connections, the stub is produced of such a length the welding

and pre- and post-heat treatments that are required in the field to make these connections, do not

cause heating of the casing to the extent the metallographic properties of the casing shells are

affected. Figure 5-62 shows a portion of a casing with the main steam connecting stubs.

Fig. 5–62 Lower Half Casing With Integral Steam Inlet Stubs

Slip ring connection

Slip ring connections are used in many high duty applications. This type of connection

does not in fact produce a joint but provides a means by which the inlet pipe can penetrate the

shell. This form of connection provides for pipe access with a flexible connection, allowing the

pipe to connect to a first stage stationary blade chamber.

Shown as Figure 5–63 is a high-pressure steam inlet connection in which the inlet pipe

penetrates the outer and inner casings to deliver steam to the inlet chamber or nozzle box using a

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combination of welded (inner pipe piece) and bolted connections to the outer casing and a slip

ring passage through the inner shell to the nozzle box.

Fig. 5–63 A High-Pressure Inlet Employing a Flexible Connection With Piston Rings

Shown in greater detail in Figure 5–64 is a portion of a slip ring system. Here the

connection is made through a guide bush with the aid of slip rings. The guide bush and slip rings

are normally produced from Stellite or a similar resistant material because of the high service in

this region and the need to minimize wear. The number of rings used depends upon the pressure

differential across the inner casing at the point of entry. The slip rings provides a precision seal,

minimizing leakage by causing successive throttlings across each ring.

Fig. 5–64 The Slip Ring System

To provide the degree of flexibility required in the inlet region, the system must provide

for relative movement between the inlet pipe, the casing, and the nozzle box or steam chamber.

To ensure this flexibility, there must be the ability for relative movement built into the ring

system. Consider the portion of the slip ring shown in Figure 5–64 where the steam enters and

flows down the pipe.

The steam, because of the pressure differential, will attempt to follow a leakage path past

the rings shown as t-t under the differential pressure dp across the rings. Each of the rings will

provides a seal surface s. There are, therefore, parallel leakage paths along each ring face as

shown. The clearance Ct allows the pipe to move in a direction transverse to its axis, and at the

same time, maintain a seal on the face s. For assembly Ct must be greater than q, which is the

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clearance between the pipe and the casing or guide bush in its normal operating position.

Clearance Ca is not as critical.

However, Ct must be sufficient to allow the ring to move and adjust inward and outward

in the radial direction without obstruction. This ensures an adequate steam seal on the inner

surface of the guide bush or casing.

The ring is shaped in a modified form such that when it is deformed to allow assembly

and released it will cause a close approximation to a circle on its outer surface. The form of such

a ring is shown in Figure 5–65. The gap g must be sufficient to enable the ring to be compressed

for assembly into the casing. Some rings are arranged so they approximate a complete circle by

arranging an overlap as shown in the detail in Figure 5–65 with gap go and gi at the outer and

inner circumferences. On compression, the ring will slide into the pipe connection and permit

assembly into the internal portion of the guide bushing or casing. Alternately, the rings may be

solid and sized to form a tight seal once located on the pipes and guide bushings and at

temperature.

Fig. 5–65 the Slip Ring

If the piston ring is used, the piston ring will not form a complete circle because it must

possess the ability to be assembled and then to close to the pipe or guide ring to affect sealing in

the vertical direction. Other seal rings are solid as shown in Figure 5–66. These rings are set to

an inner and outer diameter and by being forced together form an effective seal upon tightening

the assembly. These rings must be machined at site during assembly to achieve the correct fit.

Fig. 5–66 The Solid Ring System

5-95

Shown as Figure 5–67 is a similar inlet design and seal system, but in this case, the inlet

connection is to an intermediate pressure casing and an effective seal can be achieved using only

three rings. Figure 5–63 shows a connection for a high-pressure casing where there existed a

large pressure differential from the steam chamber to the chamber formed between the inner and

outer casings. In this design, six piston rings were used to prevent leakage. There was one more

ring preventing steam bypassing the first stage stationary blades and entering the main steam

flow without pressure expansion.

Fig. 5–67 an Intermediate Pressure Inlet System Using Slip Rings

Steam Inlet and Nozzle Box Systems

The nozzle box is discussed in chapter 6 and Figures 6–25 and 6–26show typical

arrangements of these regions of the unit.

Explosion or Relief Diaphragms

Explosion or relief diaphragms are an integral part of the low-pressure casing design and

those elements are installed on the outer upper hood. The intent of these components is to act as

a safety device to protect the blade system and the total low-pressure section against pressure

buildup and frictional overheating in the event the vacuum cannot be maintained in the unit and

steam continues to expand through the blade system.

These diaphragms are sized—singly or in combination—to pass full load steam flow to

atmospheric pressure. The pressure at which these elements are designed to rupture is set

between 5.0 and 10.0 psi above atmospheric.

These components consist of a support grating placed below a soft metal (lead or copper)

rupture diaphragm. This diaphragm is clamped firmly around its periphery to form a vacuum

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seal. While the hood pressure is below atmospheric, the diaphragm material is held in contact

with the grating as seen in Figure 5-68. However, during an emergency situation when pressure

is reversed, the diaphragm membrane is forced radially outward and it looses the support of the

grating. To ensure rupture, some device is fixed above the diaphragm to puncture it when the

deflection has reached a preset position, and upon rupture this will release the steam to

atmosphere.

Fig. 5–68 A Schematic of an Explosion or Atmospheric Relief Diaphragm

The puncturing device can be made as a circumferential knife edge or as a centrally

placed point that can be made adjustable vertically. Other rupture devises are also used and have

the capability of initiating the rupture. Various puncture systems are shown diagrammatically in

Figure 5–68. When these rupture disc membranes are replaced, it is important the replacement

part is produced from the same material and is of the same thickness because these are

parameters that affect the pressure at which rupture occurs.

High-Pressure Packing Heads

An integral part of the casings are the packing heads. These are located at each end of the

casing and are intended to provide steam seals to minimize the outward leakage of steam from

the steam path and to minimize the ingress of air that is pulled into the unit to prevent the

leakage of working fluid to atmosphere. That air is drawn into a high-pressure section requires

that some intermediate point in the total sealing system is connected to a sub-atmospheric

pressure and at this point a mixture of steam and air is carried to a special gland-sealing

condenser.

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The packing head of Figure 5–69 shows a typical configuration for a high-pressure

section. This packing head is located from the shells by bolting or in some designs may be cast

integral with the main shell portions. However, to optimize the design, the head must be able to

be aligned to the rotor to ensure the clearance at all diametral positions are at their design value.

There is an advantage to having some degree of adjustment in the packing head as this allows

finer adjustment when the unit is open for maintenance. The packing head shown in Figure 5–69

also has a secondary function of providing one face of the diffusing portion from the last stage of

the turbine section.

Fig. 5–69 The Schematic Representation of a Packing Head

The packing head is arranged to allow steam to be extracted at various axial positions to

be lead to various locations within the total thermal cycle. Some heads are also designed to have

high-pressure steam introduced into them to provide positive sealing at start-up when the unit

internal pressure would be below atmospheric.

Casing Exhaust Geometries

In the discharge region from the final stage in any turbine section, a diffuser is formed.

The shape of this diffuser is important to the total performance of the unit because it functions to

remove the steam discharging from the final rotating blade row of the section in such a manner

that there are no pressure increases formed in the downstream region from the blade. The

diffuser is formed from the walls—normally the inner from the packing head and the outer as an

attachment to the last stage diaphragm or from a portion of the casing. In Figure 5–69 was shown

a packing head with the inner diffuser surface formed from the outer conical surface of the head.

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The form of the diffuser is selected to allow the steam to expand through the passage

formed, without causing any pressure increase. In many designs, it is possible to produce a

pressure gradient through the diffuser that is decreasing from inlet to discharge. There are

significant gains to be made with correctly designed diffusers, and their form will reduce fuel

consumption by significant amounts. This is discussed in more detail in “CFD Puts New Spin on

Turbine Hoods,” in RI Fossil Plant News, Spring 1998.

Shown as Figure 5–70 is the diffuser portion from a typical reheat or low-pressure

section. Here the inner surface s-r is formed by the packing head and the outer surface t-u by

some surface provided from the major stationary components. It can be seen that there is a

gradual increase in flow area and that with a minimal or negligible increase in steam specific

volume will cause a reduction in steam velocity as it flows through the diffuser section of the

unit.

Fig. 5–70 A High-Pressure Section Diffuser, Showing the Variation of Flow Area Along

the Length of the Diffuser Portion

The steam energy expended in the low-pressure section represents a major portion of the

total energy available in the fluid, and in many unit designs, the last stage produces as much as

10% of the total power developed in the turbine. For this reason, the design parameters around

this last stage become critical to the total performance of the unit. The axial component of the

steam velocity leaving the last stage will normally be in the range of 500 to 1500 ft/sec and in

some designs may even be larger. Therefore, this indicates the criticality of the exit region of the

last stage where relatively minor changes in geometry can have a significant impact on

performance.

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To maximize the efficiency of this exhaust stage, care is taken in designing the diffuser

portion of the inner casing and also to ensure there will be as even a distributing of steam flow in

the hood as possible. The design is arranged so the pressure drop from blade exit plane to the

condenser is minimized. In addition there are often deflector vanes that are used to turn the steam

around in the upper half and divert it with minimal frictional loss into the condenser. It is normal

to employ these turning vanes and locate them to the greatest extent possible so they can act as

support bars adding rigidity to the casing structure.

Traditionally a diffuser requires a relatively long expansion passage and far more space

than is available in a conventional exhaust hood. For this reason, the designer must achieve an

acceptable form in a much shorter axial distance to minimize those losses. In addition, this

region of the exhaust hood is often produced by fabrication, which limits the geometry that can

reasonably be used.

Figure 5–71 shows the exhaust position of a long last stage blade and indicates some

forms of the diffuser plates that are quite common. The type shown in Figure 5–71a is common

in the smaller units and does provide a continuous divergence of the exhaust passage. Other

forms are as shown in figure sections b, c, and d. In some designs, the outer plate t-u is produced

as a curved form as shown in Figure 5–70. The inner plate r-s is normally straight and will often

be a portion of the bearing cone, particularly if the bearings are supported from the casing

fabrication rather than from separate pedestals mounted directly on the foundation.

This is an acceptable form, but the rate of curvature must be controlled as there is also a

tendency for premature separation of the boundary layer on this plate if the curvature is too

small. Similarly, if the outer wall is produced as shown in Figure 5–70c, then there will certainly

be flow separation at the wall discontinuity. However this form is used in some older designs.

Fig. 5–71 Various Forms of an Exhaust Stage Diffuser Arrangement

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5-101

References and Bibliography

ASME Specification A284-70a. “Low and Intermediate Tensile Strength Carbon Plate for

Machine Parts and General Construction.”

AWS, AWS Standard Welding Symbols A 2.0-68.

Blodgett, O.W. Design of Welded Structures. Cleveland, Ohio: The James F. Lincoln Arc

Welding Foundation, June 1956.

“CFD Puts New Spin on Turbine Hoods.” RI Fossil Plant News, Spring 1998, Issue 40.

Chapman, W.A.J. Workshop Technology Part III. London: Edward Arnold Ltd., 1958.

Clymer, F. “Quality Control as Applied to Steam Turbine Castings.” Intitution of Mechanical

Engineers (I. Mech. E.) Ontario Committee, 1968.

Curran, R.M. and D.P. Timo. “Heat Treated Steel for Elevated Temperature Service in Modern

Large Steam Turbines.” Presented at the Symposium on Heat Treated Steels for Elevated

Temperature Service, Mechanical Engineering Conference, Louisiana, 1964.

Fisher, J.W. Guide to Design Criteria for Bolted and Riveted Joints. 1974.

Hummer, J and J. Drahy. “The Skoda 200 MW Steam Turbine.” Czechoslovak Heavy Industry,

February 1964.

Jackson, R.L., S.A.B. Coulter, and R. Sheppard. “Importance of Matching Steam Temperatures

with Metal Temperatures During Starting of Large Steam Turbines.” Trans. American

Society of Mechanical Engineers (ASME), Vol. 79, 1957.

Kent, R.P. “Some Aspects of Metallurgical Research and Development Applied to Large Steam

Turbines.” C.A. Parsons Journal, General Electric Publication, Christmas, 1964.

Lincoln Electric Company, The Procedures Handbook of Arc Welding: Twelfth Edition, The

Lincoln Electric Company, June 1973.

Paolini, N.A. Practical Applications of Welding Technology. Canadian Welding Development

Institute.

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Peterson, R.E. Stress Concentration Factors. Wiley-Interscience Publications, New York, 1973.

Rogers, J.A. and R.C. Brewer. “Faults in Cast Components for High-pressure, High-temperature

Service.” I. Mech. E. Proc. 1964–65, Vol. 179, Pt. 1, No. 2.

Rollason, E.C. Metallurgy for Engineers. London: Edward Arnold Limited, London, 1956.

Tavernelli, T.F. and L.F. Coffin. “Experimental Support for Generalised Equation Predicting

Low Cycle Fatigue.” ASME Paper No. 61-WA-199, 1961.

Timo. D.P. “Typical Fatigue Failure Problems and Fixes in Large Steam Turbines.” A.S.M.

Conference on Fatigue, Boston 1970.

Watson, H. “Factors in the Design of Large Steam Turbines for High Availability.” Institute of

Mechanical Engineers Convention on Steam Plant Availability. Proc. I. Mech., E. 1964–5.

“What’s Ailing That Weld?” Welding Journal, August 1997.