Hydraulic Design for Energy Generation

8.1 INTRODUCTION

This chapter describes the design aspects of hydraulic structures related to the pro-duction of hydroelectric power. These structures include headrace channels; intakes;conveyance tunnels; surge tanks; penstocks; penstock manifolds; draft-tube exits; tail-tunnels, including tail-tunnel surge tanks and outlets; and tailrace channels. The pro-cedures provided in this chapter are most suitable for developing the preliminarydesigns of hydraulic structures related to the development of the hydroelectric pro-jects. To finalize designs, detailed studies must be conducted: for example, economicanalysis for the determination of penstock diameters, computer modeling of hydraulictransients for surge tank design, and studies of physical models of intake and itsapproach.

8.2 HEADRACE CHANNEL

An open-channel called the headrace channel or power channel (canal) is sometimesrequired to connect a reservoir with a power intake when the geology or topography is notsuitable for a tunnel or when an open-channel is more economical. The channel can belined or unlined, depending on the suitability of the foundation material and the projectseconomics. Friction factors for various linings used for design are as follows:

Manning’s n

Lining Minimum. Maximum

Unlined rock 0.030 0.035

Shotcrete 0.025 0.030

Formed concrete 0.012 0.016

Grassed earth 0.030 0.100

CHAPTER 8HYDRAULIC DESIGN FOR

ENERGY GENERATION

8.1

H. Wayne ColemanC. Y. Wei

James E. LindellHarza CompanyChicago, Illinois

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Source: HYDRAULIC DESIGN HANDBOOK

Headrace channels are generally designed and sized for a velocity of about 2 m/s(6.6 ft/s) at design flow conditions. Economic considerations may result in some variationfrom this velocity, depending on actual project conditions.

Channel sections are normally trapezoidal because this shape is easier to build formany different geologic conditions. The bottom width should be at least 2 m (6.6 ft) wide.Side slopes are determined according to geologic stability as follows: earth, 2H:1V or flat-ter; and rock, 1H:1V or steeper. The channel’s proportions—bottom width versus depth—are largely a matter of construction efficiency. In general, the minimum bottom widthreduces excavation, but geologic conditions may require a wider, shallower channel. Thechannel slope will result from the conveyance required to produce design velocity fordesign flow.

Channel bends should have a center-line radius of 3W to 5W or more, where W is thewater surface width of the design flow. For this radius, head loss and the rise in the watersurface at the outer bank (superelevation) will be minimal. If the radius must be reduced,the following formula can be used to estimate head loss hL:

hL � Kb�2Vg2

� (8.1)

where Kb � 2 (W/Rc), W � channel width, Rc � center-line radius, and V � mean velocity.

Superelevation will be as follows (Chow, 1959):

�Z � �2RW

c� �2

Vg

2

� (8.2)

where �Z � rise in water surface above mean flow depth.

Freeboard must include allowances for the following conditions: (1) static condi-tions with maximum reservoir level (unless closure gates are provided to isolate thechannel from the reservoir), (2) water surface rise (superelevation) caused by flowaround a curve, and (3) surge resulting from shut-off of flow downstream or suddenincrease of flow upstream. A forebay is provided at the downstream end of the head-race channel to facilitate one or more of the following: (1) low approach velocity tointake, (2) surge reduction, (3) sediment removal (desanding), or (4) storage. The fore-bay should be designed to maintain the approach flow conditions to the intake assmoothly as possible. As the minimum requirement, a small forebay should be provid-ed to facilitate good entrance conditions to the intake. It should include a smooth tran-sition to a section with a velocity not exceeding 0.5 m/s (1.64 ft/s) at the face of theintake structure

A larger forebay could be required for upsurge protection during rapid closure of tur-bine gates for load rejection. The size would be determined on the basis of the freeboardallowance for the entire headrace channel and on a hydraulic transient analysis of thechannel, if necessary.

Surge calculations should consider maximum and minimum friction factors, depend-ing on which is more critical for the case under study. Hydraulic transient (surge) studiesare generally performed using a one-dimensional, unsteady open-channel-flow simulationprogram. The computer model developed should be capable of simulating the operationof various hydraulic structures, the effect of the forebay, and operation of the power plant.Several advanced open-channel flow-simulation programs have been described by Brateret al. (1996).

8.2 Chapter Eight



HYDRAULIC DESIGN FOR ENERGY GENERATION

A large forebay is required if it will be used for diurnal storage–say, for a power peak-ing operation. In such a case, maximum and minimum operating levels would include therequired water volume, with the intake located below the minimum level. Such a forebayalso could accommodate the other three functions described above.

When the flow carries too much sediment and its removal is required to protect the tur-bines, a still larger forebay would be provided to function as a desanding basin (alsoknown as a desilting basin or desander). However, the desanding basin is more likely tobe located at the upstream end of the headrace channel. Exhibit 8.1 Illustrates a desend-ing basin. The basin can be sized using the following equation (Vanoni, 1977):

Hydraulic Design for Energy Generation 8.3

(a)

(b)

Exhibit 8.1 Sun Koshi hydroelectric project, Nepal.(a) A view of the desanding basin (looking upstream) showing concrete

guide vanes.((b) Layout Of the desanding basin.




8.4 Chapter Eight

P � (1 � e��

L

V

V

Ds�) � 100% (8.3)

where P � percentage of sediment of a particular size to be retained by the basin, L �basin length, Vs � settling (fall) velocity of suspended particles, V � mean flow velocity,and D � depth of the desanding basin. The settling velocity Vs for each particular sandparticle size can be estimated from Fig. 8.1. A separate sluicing outlet (or outlets) wouldbe provided to flush the desanding basin intermittently.

8.3 INTAKES

Most power intakes are horizontal, a few are vertical, and very few are inclined. Figures8.2, 8.3, and 8.4 are examples of the three types of intakes. Exhibit 8.2 illustrates the lay-out of a hydroelectric project with the intakes. The horizontal intake is usually connectedto a tunnel or penstock on a relatively small slope (up to 2–3 percent). The vertical intakeis frequently used in pumped-storage projects when the upper reservoir is on high ground,such as a mountain top, and a vertical shaft-tunnel is the obvious choice. An inclinedintake is used when the topography, geology, or type of dam dictate a steeper slope for thedownstream tunnel or penstock.

A variation on the three basic intake types is a tower structure, sometimes required for

FIGURE 8.1 Settling velocity as a function of particle diameter. (Dingman, 1984)





FIG

UR

E 8

.2A

typi

cal h

oriz

onta

l int

ake.

(H

arza

Eng

inee

ring

Co.

)




8.6 Chapter Eight

FIGURE 8.3 A typical vertical intake. (Harza Engineering Co.)





FIG

UR

E 8

.4A

typi

cal i

nclin

ed in

take

. (H

arza

Eng

inee

ring

Co.

)




8.8 Chapter Eight

Exhibit 8.2 (a)





Exhibit 8.2 Karun hydroelectric project, Iran(a) A vew of the dam and control structure (looking donwn-

stream showing spillway crest, radial gates, power intakes, anddiversion tunnel entrace structure.

(b) Layout of dam showing spillway, intake and powerhouse.

(b)

selective withdrawal of water. The tower includes openings with trashracks and bulkheadsat various levels, which permit water to be withdrawn from different depths to controltemperature or water quality. Computer modeling of a reservoir’s temperature and water-quality structure is generally required to finalize the required opening sites. Descriptionsof several reservoir-simulation models can be found in Brater et al. (1996). Figure 8.5 isan example of a multilevel intake tower structure for selective withdrawal.Exhibit 8.3illustrates the intake structure for a pumped storage project.

Trashracks for power intakes are designed for a velocity of about 1 m/s (3.3 ft/s) whenthe intake is accessible for cleaning. If a trashrack is not accessible for cleaning, the allow-able velocity is approximately 0.5 m/s (1.6 ft/s). Trashrack bar spacing is dictated by tur-bine protection requirements, but clear spacing of 5cm (2 in) is typical. Although head lossthrough trashracks depends heavily on the amount of clogging, the following can be usedfor a clean trashrack, (U.S. Bureau of Reclamation, 1987);




hL � Kt �2Vg

2n� (8.4)

where Vn � velocity based on the net area, Kt � 1.45 � 0.45 �AA

n

g� �

�AA

n

g�

2

, An � net area of

trashrack and support structure, and Ag � gross area of trashrack and support structure.

An intake gate is generally provided when the power tunnel or penstock is long orwhen a short penstock does not have a turbine inlet valve. This gate is provided for emer-gency closure against flow in case of runaway conditions at the turbine. The effective areaof the gate is usually about the same as that of the power tunnel or penstock, but it is rec-tangular in shape, with a height that is the same as the conduit’s diameter and a width thatis 0.8 � the conduit diameter. A bulkhead (or stop log) is provided upstream of the intakegate for servicing the gate. The trashrack slot might be used for this function by firstpulling the trashrack.

8.10 Chapter Eight

FIGURE 8.5 A typical multi-level intake tower structure for selective withdraw-al. (Harza Engineering Co.)





Exhibit 8.3 Rocky mountain pumped storage project, Georgia.(a) Intake structure of the upper reservoir.(b) Closed up view of the upper reservoir intake structure.(c) General layout of the project including upper reservoir intake, power tunnel, and power

house.

(a)

(b)




8.12 Chapter Eight

Exh

ibit

8.3

(c)





A hydraulic study is generally conducted for emergency closure of the intake gate.The maximum turbine flow or runaway flow should be considered. The runaway flowmay be 50 percent higher than the normal turbine flow for a propeller turbine. In thehydraulic study, the water levels and pressures, as well as flow into and from the gatewell, as a function of gate position are investigated (Fig. 8.6). With this information, crit-ical gate loads can be determined for the gate and hoist. The gate also may be used forpenstock filling. A minimum gate opening of 10 to 15 cm (4-6 in) is usually specifiedfor this, but a special hydraulic study must be made to determine potential gate load andvibration if the gate opens continuously by accident. In such cases, a generous gate wellor air vent must be provided downstream of the gate to provide relief once the tunnel orpenstock fills.

The head loss for a bulkhead or gate slot, including top opening, is generally about 0.1of the local velocity head at the slot. The transition length (m or ft) Lt from gate section totunnel or penstock should be approximately:

FIGURE 8.6 A typical intake gate arrangement. (Harza Engineering Co.)




8.14 Chapter Eight

Lt � �VCD� (8.5)

where V � tunnel/penstock velocity (m/s or ft/s), D � tunnel/penstock diameter (m or ft), and C � 3.00 for units in metric systems or � 9.84 for units in English sys-tems. The variation of velocity in the transition section should be as close to linear aspracticable.

Overall head loss for an intake includes trashrack, bellmouth (0.1 � V2/2g), gate slots,and transition. The potential vortex formation for an intake should be checked using Fig.8.7. Note that when the intake Froude number (V/�g�D�) exceeds 0.5, submergencerequirements increase dramatically, and the vortex formation is difficult to predict. In thiscase, a physical model study should be carried out.

8.4 TUNNELS

When the powerhouse is situated a considerable distance from the intake and when geo-logic conditions permit, a tunnel is often used to convey the flow for power generation.The size of the tunnel is dictated by economics: that is, construction cost is added to thecost of head loss (loss of generating revenue) to obtain the minimum combined cost. Thisdetermination is usually obtained by trial and error because the process does not lend itself

FIGURE 8.7 Intake submergence and vortex formation.(Gulliver and Arndt, 1991)





(a)

to a simple formula. The resulting tunnel velocity with the economic diameter is usuallyin the range of 3 to 5 m/s (10 to 17 ft/s).

The shape of the excavated tunnel normally will approximate a square bottom and a cir-cular top. The diameter of the circular top (or the width of the square bottom) should be larg-er than the required diameter. If the tunnel is lined with concrete, its cross section is likely tobe circular or have a square or trapezoidal bottom. If it is unlined or lined with shotcrete, theexcavated shape will remain, with some smoothing by filling the larger overbreak sections.

Lining is an economic consideration, balancing the cost of the lining with the powerloss caused by friction. Even an unlined tunnel will have lined sections, such as portals,and sections where rock needs extra support for geologic stability. Friction factors fordesign are as follows:

Manning’s n

Lining Minimum Maximum

Unlined 0.030 0.035

Shotcrete 0.025 0.030

Formed concrete 0.012 0.016

Minimum friction corresponds to new conditions and is used for turbine-rating andpressure-rise calculations. Maximum friction corresponds to aging and is used for eco-nomic-diameter and pressure-drop calculations. Tunnel slope is dictated by constructionsuitability and geology, with a minimum of 1:1000 for drainage during dewatered condi-




8.16 Chapter Eight

Exh

ibit

8.4

Bat

h C

ount

y pu

mpe

d st

orag

e pr

ojec

t,V

irgi

nia.

(a)

Surg

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nk o

peni

ngs

duri

ng c

onst

ruct

ion

(44-

ft in

side

dia

men

ter

and

300-

ft d

eep)

(b)

Lay

out o

f th

e pr

ojec

t inc

ludi

ng u

pper

res

ervo

r in

take

,con

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str

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re,s

urge

tank

s,po

wer

tunn

el a

nd p

ower

hous

e





tions. Tunnel bends generally have large radii for convenience of construction. Verticalbends at shafts usually have a minimum radius of 3D to minimize head loss and to pro-vide constructibility.

8.5 SURGE TANKS

Surge tanks generally are used near the downstream end of tunnels or penstocks to reducechanges in pressure caused by hydraulic transients (waterhammer) resulting from loadchanges on the turbines (ASCE, 1989; Chaudhry, 1987; Gulliver and Arndt, 1991; Moffatet al., 1990; Parmakian, 1955; Rich, 1951; Wylie and Streeter, 1993; Zipparro and Hasen,1993). A surge tank should be provided if the maximum rise in speed caused by maximumload rejection cannot be reduced to less than 60 percent of the rated speed by other prac-tical methods, such as increasing the generator’s inertia or the penstock’s diameter or bydecreasing the effective closing time of the wicket gates. In general, the provision of asurge tank should be investigated if

� 3 to 5 for units in m/s and m or

� 10 to 20 for units in ft/s and ft, (8.6)

where Li is the length of a penstock segment and Vi is the velocity for the segment(Dingman, 1984). The term �LiVi is computed from the intake to the turbine and Hn is theminimum net head.

Surge tanks normally are located as close as possible to the powerhouse for maximumeffectiveness and may be free-standing or excavated in rock. The tanks are usually vent-ed to atmosphere or can be pressurized as air chambers. The latter is not used frequentlybecause of requirements of size, air compressors, and air tightness. Exhibit 8.4 illustratesa pumped storage project with a surge tank. Figure 8.8 shows typical installations of surgetanks for controlling hydraulic transients.

Surge tanks usually are simple cylindrical vertical shafts or towers, but other geomet-ric designs are used when the surge amplitude is to be limited. For instance, an enlargedchamber can be used at the top if upsurge might cause the water level to rise above theground surface. Similarly, an enlargement or lateral tunnel or chamber is sometimes usednear the bottom of the shaft if downsurge would caused the water level to drop below thetunnel crown. When the geometry is a cylinder, analysis is relatively simple and can beperformed using design charts. If the geometry is more complicated, a hydraulic transientsimulation model is required to carry out the study (Chaudhry, 1987; Wylie and Streeter,1993; Brater et al., 1996).

Hydraulic stability for a surge tank assures that surging is limited and brief after loadchanges (Rich, 1951; Parmakian, 1955; U.S. Bureau of Reclamation 1980; Zipparro andHasen, 1993). The minimum cross-sectional area of a simple cylindrical surge tankrequired for stability can be determined using the Thoma formula:

AST � �2Agc

LH� (8.7)

where AST � minimum tank area, A � tunnel area between reservoir and surge tank, L �tunnel length between reservoir and surge tank, g � gravitational acceleration, c �

�0

0

LiVi�Hn




8.18 Chapter Eight

FIG

UR

E 8

.8A

Typi

cal v

ente

d su

rge

tank

inst

alla

tion.

Bat

h C

ount

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plan

t (19

85):

2100

MW

pum

ped

stor

age

deve

lopm

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n B

ack

Cre

ek,V

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nia.

Moo

se R

iver

pow

erpl

ant (

1987

):12

MW

dev

elop

men

t on

Moo

se R

iver

,New

Yor

k. (

Har

za E

ngin

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o.)





head loss coefficient � ��VH2� � �2

1g�

�V2

�/H2g�

�H � minimum head loss from reservoir to

surge tank, including tunnel velocity head V2/2g, and H = minimum net operating head onturbine.

For a simple surge tank (without an orifice), increase the diameter obtained from theThoma formula by 50 percent. For a typical surge tank with a restricted orifice, increasethe diameter by 25 percent. These increases are necessary to provide damping of the oscil-lation in a reasonable period of time.

Maximum upsurge in a cylindrical surge tank can be determined from Fig. 8.9. For agiven tank size, the optimum size of the orifice is based on the balanced head design sothat the maximum tunnel pressure below the surge tank equals the maximum upsurgelevel.

Maximum downsurge in a cylindrical surge tank can be determined from Fig. 8.10.Here again, the size of the orifice should be based on balanced head design as a firstattempt. However, since downsurge may differ from upsurge, and the required orifice sizemay be different for the two purposes, shaping the orifice (i.e., changing the dischargecoefficient) by rounding the top or bottom may satisfy the two area requirements approx-imately.

For maximum upsurge, use the maximum normal headwater, minimum head lossbetween reservoir and surge tank, and maximum plant flow. Assume full plant load-rejec-tion (tripout) in the shortest reasonable time.

For maximum downsurge, use the minimum normal headwater, maximum head loss,and accept load from 50 percent to 100 percent in the shortest reasonable time. At someprojects, such as pumped-storage plants, the load acceptance is criterion is more extreme;full load acceptance, is 0 percent to 100 percent in the shortest reasonable time. The con-trolling criterion will be used to design the orifice on downsurge. When the surge tankgeometry is complex (noncylindrical), a computer model should be used to determine thelimiting surge levels. (See Brater et al., 1996, for available computer models). Freeboard

FIGURE 8.8B Typical pressurized surge tank installation. Moose River powerplant (1987): 12 MWdevelopment on Moose River, New York. (Harza Engineering Co.)




8.20 Chapter Eight

FIGURE 8.9 Maximum surge in surge tank due to instantaneous stopping of flow.(Parmakian, 1955)

FIGURE 8.10 Maximum surge in a surge tank resulting from instantaneous starting of flow. (Parmakian, 1955)





for the surge tank is 10 percent of the computed rise in the water level in the surge tankfor upsurge and 15 percent of the drop in the water level for downsurge to maintain. sub-mergence of the tank invert or the orifice to avoid admitting air into the penstock.

Pressurized air chambers are often used in pumping plants for surge protection. Theyare used occasionally for power plants when the generating flow is not excessive. Thehydraulic characteristics of the chambers are complicated by the compressibility effectsof air and temperature, and the analysis does not lend itself to simple formulas and charts.A computer model is required to verify performance. Fig. 8.8(B) shows a typical airchamber design for a hydropower plant.

8.6 PENSTOCK

A penstock generally refers to a steel conduit or steel-lined tunnel connecting a reservoiror surge tank to a powerhouse (ASCE, 1989, 1993; U.S. Bureau of Reclamation, 1967;Chaudhry, 1987; Gulliver, and Arndt, 1991; Warnick et al., 1984; Wylie and Streeter,1993; Zipparro and Hasen, 1993). It is used when the internal pressure is high enough tomake a concrete-lined tunnel or unlined rock tunnel uneconomical, particularly wherecover is low.

Penstock size is usually governed by project economics. The economical diameter isdetermined by the minimum combined cost of construction and energy reduction causedby head loss in the penstock. The energy loss decreases as the diameter of the penstockincreases while construction cost increase. As with tunnels, the most economical diame-ter can be determined more accurately by a trial-and-error procedure. The following vari-ables are generally considered (U.S. Bureau Reclamation, 1967; Gulliver and Arndt,1991):

1. Cost of pipe 7. Surface roughness (friction factor)

2. Value of energy loss 8. Weight of steel penstock

3. Plant efficiency 9. Design discharge

4. Minor loss factor 10.Allowable hoop stress

5. Average head

6. Waterhammer effect

For the assessment of a preliminary design or a feasibility level, the most economicaldiameter can be estimated using the following formula (Moffat et al., 1990).

De � �CHP

0

0

.6

.4

0

3

� (8.8)

where De � the most economical penstock diameter (m or ft), H = the rated head (m orft), P = the rated capacity of the plant (kW or hp), and C = 0.52 (for metric units) or �3.07 (for English units). If the project is a small hydropower installation, the followingsimple equation can be used (Warnick et al., 1984).

De � CQ0.5 (8.9)

where De = the most economical penstock diameter (m or ft), Q = the design discharge(m3/s or ft3/sec), and C = 0.72 (for metric units) or � 0.40 (for English units).




8.22 Chapter Eight

For large hydroelectric projects with heads varying from approximately 60 m (190 ft)to 315 m (1,025 ft) and power capacities ranging from 154 MW to 730 MW, the follow-ing equation can be used (Warnick et al., 1984).

De � �Chp0.

0

6

.

3

43

� (8.10)

where De � the most economical penstock diameter (m or ft), p � the rated turbine capac-ity (kW or hp), h � the rated net head (m or ft), and C � 0.72 ( for metric units) or � 4.44(for English units).

The maximum velocity in the penstock is normally kept lower than 10 m/s (33 ft/s). Todetermine the minimum thickness of the penstock, based on the need for stiffness, corro-sion protection, and handling requirements, the following formula can be used (U.S.Bureau of Reclamation, 1967; Warnick et al., 1984).

tmin � �D

4�00

K� (8.11)

where tmin � the minimum thickness of the penstock (mm or in), D = penstock diameter(mm or in), and K = 500 (for metric units) or � 20 (for English units). After determiningthe economic diameter, check for the operating stability of the generating unit-penstockcombination using the following steps (Chaudhry, 1987; U.S. Bureau of Reclamation,1980; Warnick et al., 1984).

1. Determine the mechanical starting time in seconds for the unit Tm as

Tm � �3(6G�D2

1)0N4

2

P� (8.4)

or

Tm � �1.6(W

�R

1

2)0N

6

2

P1� (8.13)

where GD2 � flywheel effect of the turbine and generator rotating parts used inmetric system (kg-m2), WR2 � flywheel effect of the turbine and generator rotatingparts in English system (lb-ft2) � 5.932 GD2, G � weight of rotating parts (kg), D� 2 � radius of gyration of the rotating parts (m), W � weight of rotating parts(lb), R � radius of gyration of the rotating parts (ft), N � turbine speed (rpm), P �maximum turbine output (kW), and P1 � maximum turbine output (hp).

Tm is the time for torque to accelerate the rotating mass from zero to rotationalspeed. Together, the turbine runner in water, connecting shafts, and the generatordevelop the flywheel effect WR2 or GD2. The WR2 can be determined using on thefollowing formulas:

WR2turbine � 23,800 ��N

P3d/2��

5/4

(8.14)

and

WR2normalgenerator

� 356,000��KNV3/

A2��

5/4

(8.15)





where Pd � turbine rated output (hp) and kVA � generator rated output (kilovolt-amperes).

2. Determine the water column starting time for the penstock TW as follows:

Tw � (8.16)

where �(LV) � summation of product of length (measured from nearest open watersurface) and velocity for each segment of penstock from intake or surge tank to tail-race (m2/s or ft2/sec), g � gravitational acceleration (m/s2 or ft/sec2), and H � min-imum net operating head (m or ft).

3. In general, Tm/Tw2 should be maintained greater than 2 for good operating stability

and to have reasonably good responses to load changes. If Tm/Tw2 is less than 2,

there are three possible solutions:

• Increase WR2 or GD2 for the generator; this is relatively inexpensive for increas-es of up to 50.

• Increase the penstock diameter; this is probably not economical, except for anarrow range.

• Add a surge tank or move the surge tank closer to the powerhouse.

A combination of these three possible solutions may be the most cost-effective solu-tion. The following friction factors are recommended for designing steel penstocks:

�(LV)�gH

Exhibit 8.5 A typical steel penstock branch structurebeing fabricated




Penstock Age Manning’s n

New 0.012

Old 0.016

Use the value for new penstock to calculating turbine-rating and pressure-rise. To cal-culate pressure drop use the higher values.

Design pressure is determined on the basis of the turbine’s characteristics and the clo-sure rates of the wicket gates or needle valves. For Pelton turbines, closure rates are slow,and design pressure rise is usually of the order of 20 percent of the static pressure head.For Francis turbines, design pressure rise is usually 30 to 40 percent of the static pres-sure head, depending on the cost of steel lining required. A fast closure is desirable tominimize speed rise and the potential for runaway conditions in the turbine. Detailedpressure conditions are determined by a computer model that includes the water con-ductors and surge tank as well as the turbine discharge-speed characteristics and gener-ator inertia. Many computer programs capable of simulating hydraulic transients aredescribed in Wylie and Streeter, 1993. Such computer simulation studies are oftenrequired of turbine or governor manufacturers now as a part of the specifications.Ultimately, the predicted pressure conditions are verified in the load rejection tests dur-ing unit start-up.

The profile for a free-standing penstock is based on the topographic and geologic con-ditions of the ground. In other cases, the penstock may consist of shaft and tunnel sectionsthat are largely lined with concrete, with a relatively short section of steel-lined penstocknear the powerhouse.

If the penstock is free-standing, the risk of penstock rupture is greater than it is for theshaft and tunnel system. If there is a long tunnel section upstream of the free-standing pen-stock, an emergency closure valve is often added near the tunnel outlet. A hydraulic tran-sient study is necessary to determine closure conditions (by accident or because of pen-stock rupture). A vent must be provided to admit air just downstream of the valve for pen-stock rupture and must be large enough to prevent collapse of the penstock from internalsubatmospheric pressure caused by water-column separation. A free-standing penstockalso requires small air inlet-outlet valves at local high points to remove air during fillingand admit air during dewatering.

8.6.1 Penstock Branches

A penstock often delivers water to more than one turbine. In such cases, the penstock isbranched in various ways to subdivide the flow.Exhibit 8.5 illustrates a typical steel pen-stock branch structure. When the powerhouse is normal to the penstock, several configu-rations are possible (Fig. 8.11). If the powerhouse is at an angle with the penstock, a man-ifold is used (Fig. 8.12).

Head losses in branches and manifolds depend on precise geometry and often aredeveloped by model studies. However, for a typical well-designed layout, the followinghead loss coefficients can be used to estimate the head loss hb from the main into abranch:

hb � Kb�2Vg

2

� (8.17)

where V = branch velocity (m/s or ft/s); g = gravitational acceleration (m/s2 or ft/sec2); andKb = head loss coefficient 0.2 for symmetrical bifurcation, 0.3 for symmetrical trifurca-tion, and 0.2 for manifold branch.

8.24 Chapter Eight





The diameters of branched penstocks are usually determined so that the velocity isincreased significantly relative to the main penstock. Here again, the branch size is deter-mined by economics so that construction and material costs added to cost of energy lossare at a minimum. The lower limit for the size of the branch is the size of the turbine inletthat is normally provided by the turbine manufacturer. If a turbine inlet valve is provided,its diameter will either be equal to the inlet diameter or be between the inlet diameter andthe penstock branch diameter. This valve is usually a spherical type, and, as such, no headloss occurs in the fully open position. Friction losses in the branch penstocks are calcu-lated using the same friction factors used for the main penstock and the conduit lengthsup to the net head taps in the turbine inlet.

FIGURE 8.11A Example penstock branch configurations for powerhouse normal to the penstock.




8.26 Chapter Eight

FIGURE 8.11B Configurations for single bifurcated, double y-branching, and trifurcated penstocks.(Harza Engineering Co.).

8.7 DRAFT-TUBE EXITS

Draft-tubes are designed by considering the turbine’s characteristics. The net head for theturbine is based on pressure taps at the spiral-case inlet and near the draft-tube exit.Therefore, any head losses which occur after the draft-tube pressure tap are subtractedfrom the turbine net head. Because the exit head loss is generally considered to be theaverage velocity head at the end of the draft-tube, a longer draft-tube with expansion to alarger area would, in theory, reduce this loss. In actuality, however the flow is not uniform





FIGURE 8.12B Penstock manifold for an installation with six units. (Harza EngineeringCo.)

FIGURE 8.12A Examples penstock manifold configurations for a powerhouse oriented at an angle with the penstock.




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at this point; it is highly turbulent and swirling, and the true exit loss is difficult to define.Current thinking is that further extension of the draft-tube is not economical. The rule ofthumb is to end the draft-tube when the mean velocity is about 2m/s and to base the exithead loss on this velocity.

A trashrack is usually provided at the end of the draft-tube at a pumped-storage pro-ject to prevent entry of coarse debris during the pumping mode. However, during the gen-erating mode with the trashrack in place, the trashrack is subject to vibration caused bythe concentration of flow and by swirling. This complicates the design of the trashrackand increases its cost. The analysis of the rack is a combined hydraulic and structural one.The hydraulic loadings consist of drag forces on rack bars that are dependent on velocitypatterns along with pulsation of pressure caused by swirling flow. The data on hydraulicconditions can be obtained from a physical model (usually the model from the pump-tur-bine manufacturer) because fully developed mathematical models are not readily availableto predict these forces. A structural mathematical model is then applied using thehydraulic loadings obtained from the hydraulic model tests. By trial and error, thetrashrack is designed to withstand the flow-induced vibrations.

8.8 TAIL-TUNNELS

An underground power plant will have a tail-tunnel to deliver the flow to the downstreamriver or lake. For a pumped-storage project, this tunnel provides flow both ways, becauseit acts as the inlet tunnel during pumping.

For a conventional hydroelectric plant with generating only, the tunnel is usually pres-surized.. However, if the turbines are the Pelton type, the tunnel is likely to be free flowto maintain freeboard on the turbine. For a pumped-storage plant, the tunnel is most like-ly to be pressurized, because it must deliver water both ways. If the tunnel is pressurizedand is long enough, a surge chamber will be required to prevent large fluctuations of pres-sure on the turbines during load changes.

The number of tail-tunnels, usually one or two, is based on economics and con-structability. From an operational standpoint, two tunnels are desirable to allow partialoperation of the plant even during maintenance or inspection of one of the tunnels.However, two tunnels are usually more expensive than one, and usually only one will beused unless its size becomes unmanageable. The limiting size is dictated by availableequipment and tunneling methods. These factors must be evaluated carefully when esti-mating the costs of one tunnel versus two tunnels.

A manifold is used to collect the flow from the individual draft-tubes and guide theflow through a transition section to the tail-tunnel proper. This manifold is similar in con-cept to the penstock manifold, but generally the velocities are much lower. The velocityat the end of the draft-tube is typically 2 m/s (7 ft/s) and 3 m/s (10 ft/s) at the tail-tunnel.Therefore, head losses are not significant and the flow conditions are generally acceptable.A typical tail-tunnel manifold design is shown on Fig. 8.13.

8.8.1 Tail-Tunnel Surge Tanks

When an underground power plant has a significant length of pressurized tail-tunnel, asurge tank is likely to be required. The procedures for sizing and determining extremesurges are similar to the procedures used for surges in the head-tunnel, using the hydrauliccharacteristics of the tail-tunnel instead of the head-tunnel. (Refer to Sec. 8.5). Figure 8.14shows a typical tail-tunnel surge chamber.




8.8.2 Tail-Tunnel Outlet Structures

The tail-tunnel outlet structure is typically a bulkhead structure, which might incorporatesome flow spreading for energy recovery. The spreading of the flow is an economic deci-sion based on construction costs and the value of energy loss. Figure 8.15 shows a typicalstructure of a tail-tunnel outlet. If the project is the pumped-storage type, the outlet struc-ture will incorporate trashracks at the face of the structure, and the velocity at thetrashracks will be approximately 1.0 m/s (3.3 ft/s), because the racks tend to be self-clean-ing during the generating mode.

8.9 TAILRACE CHANNELS

If the outlet structure is a significant distance from the receiving waterway, a tailrace chan-nel will be required (Fig. 8.16). The sizing of the channel will be similar to that of theheadrace channel. (Refer to Sec. 82).


FIGURE 8.13 A typical tail-tunnel manifold arrangement. (Harza Engineering Co.)




8.30 Chapter Eight

FIG

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FIGURE 8.15 A typical tail-tunnel outlet structure. (Harza Engineering Co.)




8.32 Chapter Eight

FIGURE 8.16 Tailrace channels of the Guri Project. (EDELCA, Venezuela)

REFERENCES

American Society of Civil Engineer (ASCE), Civil Engineering Guidelines for Planning andDesigning Hydroelectric Developments: Vol. 2 Waterways, American Society of Civil Engineers,New York, 1989.

American Society of Civil Engineer (ASCE), Steel Penstock, ASCE Manuals and Reports onEngineering Practice No. 79, American Society of Civil Engineers, New York, 1993.

Brater, E. F., King,H. W., J. E. Lindell, and C. Y. Wei, Handbook of Hydraulics, 7th ed., McGraw-Hill, New York,1996.

Chaudhry, M. H., Applied Hydraulic Transients, 2nd ed., Van Nostrand Reinhold, New York, 1987.Chow, V. T., Open-Channel Hydraulics, McGraw-Hill, New York, 1959.Dingman, S. L., Fluvial Hydrology, W. H. Freeman, New York, 1984.Gulliver, J. S., and R. E. A. Arndt, Hydropower Engineering Handbook, McGraw-Hill, New York,

1991.Henderson, F. M., Open Channel Flow, Macmillan, New York, 1966.Moffat, A. I. B., C. Nalluri, and R. Narayanan, Hydraulic Structures, Unwin Hyman, London, UK,

1990.Parmakian, J., Waterhammer Analysis, Dover Publications, New York, 1955.Rich, G. R., Hydraulic Transients, Dover Publications, New York, 1951.U. S. Army Corps of Engineer (USACE), Hydraulic Design Criteria, U.S. Army Corps of

Engineers, Waterways Experiment Station, Vicksburg, MS, 1988.U.S. Bureau of Reclamation, Selecting Hydraulic Reaction Turbines, Engineering Monograph

No.20, Department of the Interior, 1980.





U. S. Bureau of Reclamation, Design of Small Dams, U.S. Department of the Interior, Denver, Co,1987.

U. S. Bureau of Reclamation Welded Steel Penstocks, Engineering Monograph No.3, U.S.Department of the Interior, Denver, Co, 1967.

Vanoni, V. A., ed., Sedimentation Engineering, American Society of Civil Engineers, New York1977.

Warnick, C. C., H. A. Mayo Jr., J. L. Carson, and L. H. Sheldon, Hydropower Engineering,Prentice-Hall, NJ, 1984.

Wylie, E. B., and V. L. Streeter, Fluid Transients in Systems, Prentice-Hall, Englewood Cliffs, NJ,1993.

Zipparro, V. J., and H. Hasen, Davis' Handbook of Applied Hydraulics, 4th ed., McGraw-Hill, NewYork, 1993.




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