Flange Joint

An ASME RTJ flange has a gap between two flange faces that can allow dirt to enter the fluid.

From: A Practical Guide to Piping and Valves for the Oil and Gas Industry, 2021

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Common Hidden Dangers and Remedies of Choke Lines

Sun Xiaozhen, in Common Well Control Hazards, 2013

Defect: The choke lines uses a threaded flange

Hazard

If you use a threaded joint flange, the bearing pressure capacity is lower than with the welded flange. The threaded joint flange is suitable for the low-pressure, small-diameter pipeline, however the choke lines is for the high-pressure, large-diameter pipeline (Figs. 3-2-11 to 3-2-14).

Fig. 3-2-11. Threaded joint flange sectional drawing.

Fig. 3-2-12. Threaded joint flange (A).

Fig. 3-2-13. Threaded joint flange (B).

Fig. 3-2-14. Threaded joint flange (C).

Remedy

There are two methods to connect the flange: one is the threaded joint flange and the other is the welded flange. The relief line belongs to the high-pressure, large-diameter pipeline; as a result, it should use the welded flange instead of the screw joint flange (Figs. 3-2-15 to 3-2-18).

Fig. 3-2-15. Flat welding flange.

Fig. 3-2-16. Friction welding flange with neck.

Fig. 3-2-17. Welded flange (A).

Fig. 3-2-18. Welded flange (B).

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Friction clutch

Heinz Heisler MSc., BSc., F.I.M.I., M.S.O.E., M.I.R.T.E., M.C.I.T., M.I.L.T., in Advanced Vehicle Technology (Second Edition), 2002

2.5 Clutch misalignment (Fig. 2.10(a–d))

Clutch faults can sometimes be traced to misalignment of the crankshaft to flywheel flange joint, flywheel housing and bell housing. Therefore, if misalignment exists, the driven plate plane of rotation will always be slightly skewed to that of the restrained hub which is made to rotate about the spigot shaft's axis. Misalignment is generally responsible for the following faults:

1

Rapid wear on the splines of the clutch driven plate hub, this being caused mainly by the tilted hub applying uneven pressure over the interface length of the splines.

2

The driven plate breaking away from the splined hub due to the continuous cyclic flexing of the plate relative to its hub.

3

Excessively worn pressure plate release mechanism, causing rough and uneven clutch engagement.

4

Fierce chattering or dragging clutch resulting in difficult gear changing.

If excessive clutch drag, backlash and poor changes are evident and the faults cannot be corrected, then the only remedy is to remove both gearbox and clutch assembly so that the flywheel housing alignment can be assessed (Fig. 2.10).

Fig. 2.10(a–d). Crankshaft flywheel and clutch housing alignment

2.5.1 Crankshaft end float (Fig. 2.10(a))

Before carrying out engine crankshaft, flywheel or flywheel housing misalignment tests, ensure that the crankshaft end float is within limits. (Otherwise inaccurate run-out readings may be observed.)

To measure the crankshaft end float, mount the magnetic dial gauge base to the back of the flywheel housing and position the indicator pointer against the crankshaft flanged end face. Zero the dial gauge and with the assistance of a suitable lever, force the crankshaft back and forth and, at the same time, observe the reading. Acceptable end float values are normally between 0.08 and 0.30 mm.

2.5.2 Crankshaft flywheel flange runout (Fig. 2.10(a))

The crankshaft flange flywheel joint face must be perpendicular to its axis of rotation with no permissible runout. To check for any misalignment, keep the dial gauge assembly mounted as for the end float check. Zero gauge the dial and rotate the crankshaft by hand for one complete revolution whilst observing any dial movement. Investigate further if runout exists.

2.5.3 Flywheel friction face and rim face runout (Fig. 2.10(a and b))

When the flywheel is centred by the crankshaft axis, it is essential that the flywheel friction face and rim rotate perpendicularly to the crankshaft axis.

Mount the dial gauge magnetic base to the engine flywheel housing. First set the indicator pointer against the friction face of the flywheel near the outer edge (Fig. 2.10(a and b)) and set gauge to zero. Turn the flywheel one revolution and observe the amount of variation. Secondly reset indicator pointer against the flywheel rim and repeat the test procedure (Fig. 2.10(b)). Maximum permissible runout in both tests is 0.02 mm per 20mm of flywheel radius. Thus with a 300mm diameter clutch fitted, maximum run-out would be 0.15 mm. Repeat both tests 2 or 3 times and compare readings to eliminate test error.

2.5.4 Flywheel housing runout (Fig. 2.10(c))

When the gearbox bell housing is centred by the inside diameter and rear face of the engine flywheel housing, it is essential that the inside diameter and rear face of the housing should be concentric and parallel respectively with the flywheel.

Mount the dial gauge magnetic base to the flywheel friction face and position. Set the indicator pointer against the face of the housing. Make sure that the pointer is not in the path of the fixing holes in the housing face or else incorrect readings may result. Zero the indicator and observe the reading whilst the crankshaft is rotated one complete revolution. Reset the indicator pointer against the internally machined recess of the clutch housing and repeat the test procedure. Maximum permissible runout is 0.20mm. Repeat both tests two or three times and compare readings to eliminate errors.

2.5.5 Detachable bell housing runout (Fig. 2.10(c and d))

When the gearbox bell housing is located by dowel pins instead of the inside diameter of the engine flywheel housing (Fig. 2.10(c)) (shouldered bell housing), it is advisable to separate the clutch bell housing from the gearbox and mount it to the flywheel housing for a concentric check.

Mount the dial gauge magnetic base onto the flywheel friction face and position the indicator pointer against the internal recess of the bell housing gearbox joint bore (Fig. 2.10(d)). Set the gauge to zero and turn the crankshaft by hand one complete revolution. At the same time, observe the dial gauge reading.

Maximum permissible runout should not exceed 0.25mm.

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Piping Components

Peter Smith, in The Fundamentals of Piping Design, 2007

Dimensional Standards for ASME Flanges

The dimensional standards for the various types of flanges just mentioned are covered in two ASME standards: ASME B16.5, NPS ½″ through 24″ (DN 15 through 600) and ASME B16.47, NPS 26″ through 60″ (DN 650 through 1500). Depending on the jointing method between the flange and the pipe, one of the following ASME dimensional standards will apply: ASME B16.25 for butt-weld ends, ASME B1.20.1 for threaded ends, or ASME B16.11 for socket-weld ends.

Standard B16.5, 2003, Pipe Flanges and Flanged Fittings: NPS ½″ through 24″, covers pressure-temperature ratings, materials, dimensions, tolerances, marking, testing, and methods of designating openings for pipe flanges and flanged fittings. Included are flanges with rating class designations 150, 300, 400, 600, 900, 1500, and 2500 in sizes NPS ½ through NPS 24, with requirements given in both metric and U.S. customary units with diameter of bolts and flange bolt holes expressed in inch units; flanged fittings with rating class designation 150 and 300 in sizes NPS ½ through NPS 24, with requirements given in both metric and U.S. customary units, with diameter of bolts and flange bolt holes expressed in inch units; and flanged fittings with rating class designation 400, 600, 900, 1500, and 2500 in sizes NPS ½ through NPS 24 that are acknowledged in Annex G, in which only U.S. customary units are provided.

This standard is limited to flanges and flanged fittings made from cast or forged materials, blind flanges, and certain reducing flanges made from cast, forged, or plate materials.

Also included in this standard are requirements and recommendations regarding flange bolting, flange gaskets, and flange joints. The subject matter is as follows:

Committee Roster.

Correspondence with the B16 Committee.

(1)

Scope.

(2)

Pressure-Temperature Ratings.

(3)

Component Size.

(4)

Marking.

(5)

Materials.

(6)

Dimensions.

(7)

Tolerances.

(8)

Pressure Testing.

Standard B16.47, Large Diameter Steel Flanges, covers pressure-temperature ratings, materials, dimensions, tolerances, marking, and testing for pipe flanges in sizes NPS 26 through NPS 60 and in ratings classes 75, 150, 300, 400, 600, and 900.

Flanges may be cast, forged, or plate (for blind flanges only) materials, as listed in Table 1A. Requirements and recommendations regarding bolting and gaskets are included. The subject matter is as follows:

Standards Committee Roster.

(1)

Scope.

(2)

Pressure-Temperature Ratings.

(3)

Size.

(4)

Marking.

(5)

Materials.

(6)

Dimensions.

(7)

Tolerances.

(8)

Test.

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Microstructure and elemental features of stainless-steel surface

Nagamitsu Yoshimura, in A Review: Ultrahigh-Vacuum Technology for Electron Microscopes, 2020

3 Vacuum characteristics

A vacuum fired (1050°C) SS304 system was manufactured using the procedure described next.

Vacuum brazing (Ni-based brazing alloy, 1050°C) was applied to the chamber joints with an exception of Conflat flange joints which were attached by argon-arc welding. Consequently, most of the chamber walls were vacuum fired (1050°C, 30 min) in a vacuum furnace. The density of H2 molecules dissolved in the interior of SS304 walls of 4 mm thickness should be much reduced[3,16] during the vacuum firing period.

A SIP with Ti-sublimation pump filaments (conventional ones) was newly designed and manufactured for XHV pumping. The nominal pumping speed of the SIP was 75 L s−1 for N2, and the N2 pumping speed of the Ti-sublimation pump (just after fresh getter films were deposited) was estimated 400 L s−1. The (d×B) product value (d, anode cylinder diameter; B, magnetic flux) of the Penning discharge cells was designed as high as 3.1 kG cm on the average for effective pumping in the ultrahigh vacuum range. This value (3.1 kG cm) of (d×B) product is much higher than those of commercial SIPs. As Rutherford discussed,[26] a high value of (d×B) product of the Penning cells is very effective for XHV. Indeed, the SIP has a considerably high measured pumping speed in the 10−6–10−7 Pa range. For instance, the pumping speeds for N2 at 10−6 and 10−7 Pa were measured to be 70 and 45 L s−1, respectively.

The chamber was equipped with a conventional quadrupole MS (tungsten filament) in addition to a Bayard–Alpert ionization gauge (BAG, tungsten filament, nude type) and an extractor ionization gauge (EXG, iridium filament coated with thorium oxide, nude type). Eleven Conflat-type flanges (mostly 70 mm diam, blind type) were additionally equipped to the chamber. The SIP was connected to the chamber directly, and a conventional TMP was connected through a conventional all-metal valve. The volume and surface area of the chamber (including those of the SIP vessel and MS head) were approximately 8 L and 4000 cm2, respectively.

Pressures of 10−8 Pa range were indicated by the BAG and EXG in the chamber which had been in situ baked (about 300°C–1 day and 200°C—1 day). The residual gas spectrum is typical, in which the H2 ion intensity was the highest. And it was found that the major source was the MS head itself with an incandescent tungsten filament. It was also found that the BAG indicated a pressure (N2 equiv) higher than the pressure indicated by the EG. The pressure indication by the BAG was not lower than 10−8 Pa even when the EG indicated a low 10−9 Pa. The pressure indication by the EXG is believed to be more reliable, as Beeck and Reich[27] discussed.

The pumping characteristics were measured by the EG in a long pumping time after the chamber (including the SIP vessel) was exposed to the moist atmosphere for 15 min. The pressure (N2 equiv) in the chamber with the MS being off was approximately 2×109 Pa just before the air exposure, 10 days after the in situ bakeout (200°C and 300°C).

After 2 days’ evacuation by the TMP following the air exposure, the pressure indicated by the EG was 9×107 Pa. Then, the chamber (including the SIP vessel and MS head) was in situ baked at 80°C for 20 h. After 60 h evacuation by the SIP with the metal valve closed, the pressure was 5.4×108 Pa. And, the chamber (including the pump vessel and MS head) was further in situ baked at 170°C for 1 day. The EG was degassed with electron bombardment (19 W, 3 min), and three Ti filaments were degassed (40 A) for 4 min each. The Ti getter was flashed (50 A, 90 s each) three times in the 10−9 range. The pressure just before flashing Ti getter was about 2.1×109 Pa. On the next day an XHV of 1.5×109 Pa was indicated by the EG with the MS and BAG being off, as an ultimate pressure. The pressure would not significantly change with elapsed pumping days, though the pressure varied a little around 1.5×109 Pa with the variation of the room temperature.

The dominant residual gas molecules must be H2. It should be noted that the indicated pressure of 1.5×109 Pa was a N2 equiv pressure.

In order to measure the outgassing rate of the chamber wall, an oxygen-free copper disk plate (121 mm diam) with an aperture of 3.3 mm diam was used as a gasket for the inlet flange of the SIP. The pumping conductance of the orifice of 3.3 mm diam was calculated as 1.0 L s−1 for N2 and 3.7 L s−1 for H2 at the room temperature. The MS head and BAG were removed, and blind Conflat flange (70 mm diam) was installed. Thus the surface area of the chamber walls, which separated the orifice plate, was reduced to 2500 cm2.

The chamber with the SIP (kept off) was evacuated by the TMP through the metal valve following reassembling the chamber system. After 17 h evacuation, the pressure indicated by the EG was 1.6×106 Pa. Then, the chamber and the SIP vessel were in situ baked at 170°C for 1 day. The SIP was switched on soon after the baking heater was switched off. In an interval of 1 h after cessation of bakeout, the following treatments were sequentially conducted: three Ti filaments of the Ti sublimation pump were sequentially degassed (40 A, 3 min each), and the Ti getter was flashed (50 A, 90 s each) three times, in sequence. Then, the EG was degassed with electron bombardment (19 W, 3 min). The metal valve was closed at about 4.0×106 Pa, 1 h after cessation of bakeout. The pressure jumped up to about 5.0×105 Pa when the valve was closed, and it went down smoothly into the 10−8 Pa range. The chamber walls were still hot when the metal valve was closed. The pump-down characteristics, after the metal valve closed, were measured by the EG. The pressure indicated on the next day was about 5.0×108 Pa.

We can estimate the outgassing rate of the chamber wall after an in situ bakeout (170°C, 1 day) as a N2 equiv value. The outgassing rate as a N2 equiv value, as a function of the elapsed time t after the metal valve closed, was roughly estimated from the pressure P(t) (N2 equiv) as K (N2 equiv)=P(t) (N2 equiv) (Pa)×1(L s−1)/2500 (cm2), by neglecting the outgassing of the copper disk plate (121 mm diam).

The characteristics of the outgassing rate K (N2 equiv) thus calculated are presented in Fig. 3. The outgassing rate K (N2 equiv) at the elapsed time of 15 h was estimated as 2.2×1011 Pa L (s cm2)−1, by neglecting the outgassing of the copper disk plate (121 mm diam).

Fig. 3. Outgassing rate K (N2 equiv) of the vacuum-fired SS304 chamber wall, as a function of the elapsed time after the metal valve closed. The rate was calculated from the pressure in the chamber evacuated through an orifice (3.3 mm diam 1 L s−1 for N2). The pressure was measured by the EG as N2 equiv pressure.

It was natural to assume that the species of residual gas molecules was mostly H2. The outgassing rate for the actual gas species must be considerably higher than the calculated outgassing rate as a N2 equiv value.

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Start-up Sequence and Commissioning Procedures

Alireza Bahadori Ph.D., in Natural Gas Processing, 2014

16.10.1 Tightness test of sections operating under vacuum

1.

When sections of units are normally operated under vacuum conditions, it is necessary to test for proper tightness.

Being almost impossible to find leaks while the section is under vacuum, the most common way to do the test is to check all flanges, joints, valves packing, pumps seal, and any other possible point of leaks by pressurizing the system up to 80% of the lower design pressure of equipment included in the system to be tested or 80% of pressure relieve valves installed. Whichever is the lower will govern the pressure test.

2.

Once the system or section is at the selected pressure conditions, all joints shall be carefully checked with soapy solution, or equivalent, and even minimum leaks shall be repaired. Then, the vacuum shall be pulled at the selected value. Stop vacuum pulling devices and check for pressure drop in the selected period of time (normally 1 h).

If the pressure drop is higher than specified, the system shall be pressurized again and rechecked to find the leaks.

Repair possible leaks and pull the vacuum again. The operation shall be repeated until the pressure drop is within given limits.

3.

When the system to be vacuum-tested is very large, it may be convenient to split it in two or more sections by closing isolating valves. Installation of blinds is not recommended.

4.

Only when the system does not show any leak from the joints, possible internal leak due to bad-seal of valves may be suspected. In this case, it is convenient to blind-off suspect valves.

When suspect bad-sealing valves are the welded type, it will be necessary to open their cover to inspect the internal disc, and/or seat for proper smoothness and/or full travel of the wedge into the ring seats.

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CASING AND PIPELINE CORROSION

George V. Chilingar, ... Ghazi D. Al-Qahtani, in The Fundamentals of Corrosion and Scaling for Petroleum & Environmental Engineers, 2008

5.4.1 Wellhead Insulation

Use of electrical insulation stops current flow down the casing from the surface and reduces both internal and external casing corrosion. Dielectric insulation materials for both screw and flange joints are commonly used to insulate casing from flowlines. Insulation of wells by connecting them to a single battery is often recommended. It should be noted that when the flowline is at high potential due to cathodic protection, it may induce interference corrosion. In this case, the insulating joints may be partially shunted or wellhead potential elevated by attaching a sacrificial anode (Figure 5.4). Heat-resistant material should be selected for hot, high-pressure wells to prevent failure of insulation materials (also see Jones, 1988, p. 64).

Figure 5.4. Installation of galvanic anodes.

(after NACE, Houston, TX, Control of Pipeline Corrosion, 1979, figure 8-6)Copyright © 1979
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FAILURE OF AUSTENITIC STAINLESS STEEL COMPONENTS USED IN NITROGEN OXIDE PLANT

V.M.J. SHARMA, ... T.S. LAKSHMANAN, in Failure Analysis Case Studies II, 2001

Abstract

Austenitic stainless steel components of a nitrogen oxide plant have been found to leak in service. The failed components, namely pipe-to-pipe joints and pipe-to-flange joints, have been studied through standard metallographic techniques to analyse the cause of the failure. In case study I, involving the failure of pipe-to-flange joints, cracking was observed in the pipe wall next to the pipe-to-flange weld. In case study II, involving the failure of a sight port flange, cracking was observed in the flange adjacent to the pipe-to-flange weld. In both cases, cracking was by an intergranular mechanism, and carbon contents were much higher than permitted for “L” grades of austenitic stainless steel. © 1997 Elsevier Science Ltd

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Fatigue Design of Components

G R Razmjoo, P J Tubby, in European Structural Integrity Society, 1997

6 CONCLUSIONS

Testing arrangements were developed to allow tests to be conducted under biaxial bending and rotating principal stress conditions for non-load carrying fillet welded joints as well as load carrying fillet welded tube-to-flange joints in steel. Preliminary results allow the following tentative conclusions:

i)

Under applied equibiaxial bending condition the biaxiality ratio adjacent to a welded joint is strongly dependent on joint geometry. Typical values measured in this programme were 0.39 to 0.45 for circular cover plates and 0.62 to 0.78 for square cover plates.

ii)

Cover plates under biaxial bending gave lower fatigue endurances than those under uniaxial loading when expressed in terms of principal stress range.

iii)

Out-of-phase torsion and tension loading can reduce the fatigue life of fillet welded joints in comparison to that under in-phase loading. A life reduction of the order of magnitude life reduction was noted when expressed in terms of principal stress range.

iv)

The use of maximum principal stress range as the fatigue damage parameter resulted in satisfactory fatigue life prediction for the case of in-phase loading but showed that it could be unsafe for the case of out-of-phase tension and torsion loading.

v)

The fatigue design methods in the current design code1 may not adequately predict the fatigue life under the condition of rotating principal stress.

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Common Hidden Dangers and Remedies for Choke and Kill Manifolds

Sun Xiaozhen, in Common Well Control Hazards, 2013

Hazard

At present, almost all Chinese on-land drilling uses steel pipeline as the relief line. The choke line and the choke manifold and throttle are connected through a flange, and the axis deviation of the flange joint should be very small, otherwise it is difficult to buckle a steel ring into the flange groove. Using a nonadjustable height throttle manifold pier base, in order to make sure the relief line is connected to the throttle and well killing manifolds, we can only fill rock and soil under the pier base (Figs. 4-1-13, 4-1-14) to adjust the throttle manifold height. This is inconvenient and time-consuming; moreover, the base foundation is not firm. In Figures 4-1-15 and 4-1-16, the manifold pier base height is not adjustable and the installation is very difficult.

Fig. 4-1-13. Fill stones or wood bricks under the throttle manifold pier base.

Fig. 4-1-14. Fill loose sand under the throttle manifold pier base.

Fig. 4-1-15. Nonadjustable pier base height (A).

Fig. 4-1-16. Nonadjustable pier base height (B).

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Design and Applications

R.J. Lewandowski, W.F. BrittJr., in Comprehensive Composite Materials, 2000

6.31.3.2.6 Design approach

Contact molded FRP pipe, made according to PS 15-69 (1969), is normally rated by pressure in increments of 25 psi up to 150 psi. Standard machine made pipe does not follow this type of rating and the designers should refer to the ultimate pressure rating listed in the manufacturer's catalog. It is important to note that these ratings are based on the allowable pressure of a continuously supported pipe subject to pressure stress only. Since piping systems are almost never continuously supported, the stresses of bending must be considered when determining the allowable working pressure of an FRP pipe. The equations presented in Section 6.31.3.3 take these stresses into account.

Wall thickness is based on a 10:1 safety factor in the hoop direction and it is customary to use a 5:1 or 6:1 safely factor in the axial direction. By maintaining these safety factors in the design, stress risers in elbows and other fittings will be within the allowable stress limit.

The method of design layout preferred by many engineering firms is an anchor to anchor design. This design method can be economical and offers many advantages. The anchor to anchor system is more rigid and less susceptible to damage due to dynamic loading. This system also provides a means for controlling expansion (thermal and pressure), thereby reducing the length of offsets and eliminating the need for expansion joints. Anchors are placed on either side of every change in pipe direction and as near to the fitting as possible. The amount of structural steel required to absorb the loads imposed on the anchors (see Figure 4) can be minimized by keeping the pipe elevation close to the steel and by utilizing tension members between anchors. It is important to recognize that the pipe must be guided between anchors to prevent buckling. In cases where there are long straight runs, anchors should be placed no more than 150–200 feet apart.

Fig. 4. Anchor loads.

In many cases where the anchor to anchor design is not used, the pipe is often simply hung with rod hangers (see Figures 5–7). This is an example of a highly flexible piping system. Expansion is uncontrolled and dynamic forces can cause very large movements of the pipe system. This design will work where the operating temperature is near ambient temperature and fluid velocities are very low. However, even mild fluid dynamic forces can destroy a pumped system that is installed in this manner. Vibrations induced by the pump can damage sections of the system where frequencies are within the resonant range. Wind loads can also induce damaging stresses. To reduce vibration and wind load effects, the pipe should be laterally restrained at specified intervals along the pipe. These restraints should not be located near changes in direction where offset legs are required for flexibility.

Fig. 5. Support spans.

Fig. 6. Offsets.

Fig. 7. Guide spacing.

Many piping system failures that occur during hydrotesting, or even after years of service, have been attributed to poor workmanship. In reality, failures are caused by overstress and this could be due to poor workmanship and/or unsatisfactory design. The task of the designer is to eliminate, as much as possible, the likelihood of failure. With sufficient effort spent on the design and in the instruction of installation personnel, failures can be effecreduced.

The design techniques presented here have helped to standardize a conservative design approach. Beyond this, a great deal of work is necessary to standardize installation techniques, especially the methods for fabricating joints.

In an effort to reduce the possibility of failure, the designer should seriously consider the following:

(i)

Keep the pipe run away from high traffic areas where damage from equipment impact is likely.

(ii)

Keep flange joints to a minimum. Flanges are expensive components and sources of leaks.

(iii)

Provide vents at each high point to allow air to be removed from the system prior to testing and system start up.

(iv)

Provide drains at each low point or pocket. Drains with blind flanges will allow the line to be drained if repairs are necessary.

(v)

Ensure that all supports, anchors, and guides are installed prior to the hydrotest. This cannot be overemphasized since the pipe system can be severely damaged without proper pipe support.

(vi)

All valves and valve operators or components in the system must be independently supported.

(vii)

Valves that require high torque to open and close should be anchored so that the high torque does not damage the pipe.

(viii)

Riser supports for vertical runs should be guided or laterally restrained to reduce vibration and effects of wind load.

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