More Questions on bolt torque 2

[ProcEng2]

Is there an ASME Code Interpretation that states how high the preload can be as a percentage of yield strength due to bolt torque?

I ask this question after reading thread 292-95664 because in the application we are considering, the bearing
stress under the head could exceed 80% of Sy at the max of
requested 268-387 ftlb torque for a SA-193 Gr. B7 3/4-10 B18.3 fastener in a 7/8 in. hole, without a hardened washer...

I would think if the bearing stress exceeded Sy, there would be a concern for loss of preload. This is a
critical application for a vertical pump pedestal for vibration,we can not afford to loose pre load at the
pedestal to sole plate.

 

[BigInch]

You're using A-193's for anchor bolts?

BigInch-born in the trenches.

http://virtualpipeline.spaces.msn.com

 

[ProcEng2]

This is a OEM vertical pump application with a right angle gear drive mounted on a pedestal, the pedestal, bolted to a sole plate, the sole plate grouted to the foundation (and bolted). All of it, including the pedestal mounting bolts, are OEM supply/design. The issue is that the pedestal has notched ribbes in four places that are very close to the heavy hex B7 bolts such that it makes it difficult to get a torque wrench on 4 of the 8 bolts. The problem being that torque of these bolt for vibration free operation is critical because the vertical pump is long. Someone felt that the replacement of the B7 bolts with Hex socket ANSI B18.3 capscrews would make it easier to torque. The two problems I have with this is 1) Standard A-574 material is only in ASME III by code case because it is very strong and can be brittle. B7, I would like to stay with due to notch toughness/ductility. 2) The head on B18.3 capscrews provides significantly less bearing area, and the hole is 7/8 in. further reduces. This will require addition of a washer and review of the thread engagement available in the sole plate. But the only criteria I can find is 2Sm for Class 1 bolting, my application is Class 3. 2 Sm is 70 ksi at 100 F. I estimate that the bearing stress will be 86 ksi using min. area without a washer, 77 ksi with a washer. So my question is do I need to use 2 Sm or is there a Code Interpretation that applies just to preload. I had always thought once a joint was loaded, other loads had a tendency to reverse or unload the joint...

Thanks for the reply, but it looks to me on the surface to be an over stress condition and not acceptable.

 

[arto]

Look in ASME PCC-1 they give preloads as 50ksi on B7's

 

[LSThill]

Subject: VI. Connection Set Up, Torque Calculations, and Meridium Population
1. Identify and generate all applicable connections in Meridium for each exchanger.
2. Perform bolt torque calculations to establish bolt torque limits for each flange.
3. Populate Meridium with the applicable information for each connection
4. WRC 408. WRC 505, WRC 507,WRC 508,WRC 510; AND Axi/PRO and Axi/PRO Plus!


Axi/PRO and Axi/PRO Plus!


Axi/PRO provides FEA calculations for Flanges and Other Axisymmetric Models.

Bolts, nuts and holes may be included in the 3D brick models so that users can see the results of hole spacing on the stress distributions. Dimensionally accurate flange models including studs and nuts are generated automatically for six major flange standards.


Build 3D models for the following code evaluations:
• ASME Section VIII, Div. 1, Appendix 2
• ASME BFJ
• EN 13445


Newsworthy:

European Union IPPC directive 96/61/EU

Mandates that all processing plants and power stations improve the management of industrial processes and ensure a higher level of protection for the environment. An important part of this legislation is the reduction of fugitive emissions.

To put the fugitive emissions problem in perspective, a typical European refinery loses between 600 and 10,000 tons of volatile organic compounds (VOCs) per year. It is estimated that 72% of this is attributable to plant equipment such as pipe flanges, pumps, valves and vessels. And one major chemical manufacturer reports that 28% of their emissions come from flanges.

It is a fact that all gaskets leak, but do you know the variance between gasket types? Most gasket manufacturers cannot answer this question. Selection of the correct gasket is critical. Process plant managers need to know how badly their gaskets leak. The standard practice of gasket selection based on fluid, temperature, pressure and bolting is no longer acceptable.

Consider the actual costs of not knowing:
• Loss of Product
• Loss of Production Time
• Clean-up Costs
• Labor Cost to Repair Leaks
• Environmental Fines
• Bad Community Relations
Axi/PRO, flange analysis software that not only protects the environment but saves process plants money.



BOLTED FLANGE ASSEMBLY: PRELIMINARY ELASTIC INTERACTION DATA AND IMPROVED BOLT-UP PROCEDURES
G. BIBEL AND R. EZELL

WRC Bulletin 408
January 1996 (27pp)(ISBN#1-58145-407-4)
Design bolt loading in a bolted flange connection is difficult to obtain because of elastic interaction. Elastic interaction data is presented for two different pipe flanges, three different gasket styles and two different bolting patterns. An algorithm that compensates for elastic interaction is given. Uniform bolt preload (±2%) is obtained in a single pass bolt-up operation. The algorithm is based on experimentally obtained interaction coefficients.
Important results are:
• Elastic interaction can cause individual bolts in flanges with spiral wound gaskets to lose up to 98% of their initial preload when adjacent bolts are tightened.
• The average bolt stress can be 25-50% below design with a spiral would gasket even with a multi-pass (3 pass) bolting procedure.
Publication of this document - WRC Bulletin No. 408 was sponsored by the Committee on Bolted Flanged Connection of the Pressure Vessel Research Council.


AN OVERVIEW AND VALIDATION OF THE FITNESS-FOR-SERVICE ASSESSMENT PROCEDURES FOR LOCAL THIN AREAS IN API 579
Janelle, J., Osage, D. A., Burkhart, S. J.

WRC Bulletin 505
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September 2005 ISBN: 1-58145- 512-7 231 pages
Most of the US design codes and standards for pressure containing equipment do not adequately address degradation and damage during operation. In the pressure vessel and pipeline industries, surface flaws are major limiting factors of vessel or pipe remaining life, and this type of degradation due to age and aggressive environment eventually threatens the structural integrity of equipment. Replacing vessel and piping equipment is expensive, making it cost effective and desirable to operate damaged pressurized equipment. For corrosion beyond a specified limit or other damage mechanism like cracking, a Fitness-For-Service (FFS) assessment is required. The specific objectives reported in the bulletin are:
• Objective 1 – Validate the API 579 Section 5 LTA rules in addition to the validation in WRC 465. Significant validation work was reported in WRC 465. After publication, additional burst pressure results became available. The validation now includes comparison of the API 579 methodology to other industry methods and to a database of full scale test results.
• Objective 2 – Develop new or improve upon the existing methodology to increase the accuracy of the assessment procedures and eliminate some of the limitations. There are twenty-five different methods compiled in this study for analyzing local thin areas in pipes and vessels. These analysis methods all have roots in various industries, codes, and standards. In industry, at least five of these methods are actively used in fitness-for-service assessments today. This can make communication difficult between parties using different assessment procedures, and some parties may be using methods with low accuracy or reliability. Depending on the assessment code that is used, assessment results may vary drastically. One standardized set of analysis guidelines is needed to eliminate confusion regarding which method should be used
• Objective 3 – Standardize the in-service margin between MAWP and failure pressure for industry analysis methods and different construction code margins on allowable stress ¬The LTA assessment methods in API 579 may be used to evaluate equipment built to many different construction codes. Because the design margins in these codes are different, a method needs to be introduced that permits scaling of LTA assessment results such that a consistent in-service margin can be obtained based on the ratio of predicted burst pressure to the maximum allowable working pressure computed using the appropriate procedures in each construction code.
• Objective 4 – Improve the existing rules for LTAs subject to supplemental loading (circumferential extent of the LTA). Significant improvements have been made to the original method contained in API 579 published in 2000.
• Objective 5 – Introduce assessment procedures for the evaluation of HIC damage. An assessment method based on a remaining strength factor has been developed for sub-surface and surface breaking HIC damage.
• Objective 6 – Introduce procedures for the evaluation of local metal loss in cylindrical shells subject to external pressure. A simplified method has been developed to evaluate metal loss in cylindrical shells subject to external pressure using existing construction code rules for external pressure evaluation.


LONG TERM PRESSURIZED GRAPHITE GASKETED JOINT TESTS
Luc Marchand, Michel Derenne, Olivier Sakr, Hakim A. Bouzid

WRC Bulletin 507
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Report 1: Long Term Pressurized Graphite Gasketed Joint Tests
Luc Marchand, Michel Derenne, Olivier Sakr, Hakim A. Bouzid

Report 2: Long Term Performance Of Flexible Graphite Based Gaskets
Under Nitrogen Exposure; Luc Marchand, Michel Derenne

Report 3: Additional Long Term Pressurized Joint Tests On
Corrugated Flexible Graphite Gaskets; Luc Marchand, Michel Derenne

The long term performance and elevated temperature capabilities of flexible graphite based gaskets have been the focus of several PVRC research projects. This WRC Bulletin combines a series of recent, inter-related research projects which have been commissioned to further quantify the long term performance of flexible graphite gaskets in different atmospheres, and to validate earlier service life and performance predictions which are based on shorter duration exposure.

Initiated in 1996, PVRC project 96-09G was the first of a series of pressurized media tests launched to investigate the effect of long term pressurized hot air exposure on the performance of two different types of reinforced flexible graphite based gaskets that are commonly used in industry today. One a foil reinforced sheet type flexible graphite gasket material and the second a corrugated insert flexible graphite gasket. The project was designed to supplement and confirm the long term graphite gasket life predictions that are based upon the results of previous PVRC projects 91-8, 93-3 and 94-25. While the long duration stress relaxation and leak rate results correlated well with the quantification methods identified in the earlier research for the reinforced sheet type material, the resultant performance of the corrugated insert type gasket was not as expected and, at the time, could not be explained.

PVRC project 00-BFC-12G was commissioned to investigate the possible causes for the fast rate of degradation that occurred with the corrugated insert flexible graphite gaskets in the earlier PVRC 96-09G investigation. This experimental work studied the effect of long term pressurized hot air exposure on the performance of two different brands of corrugated insert flexible graphite gaskets, and in particular identified the effect of three important parameters on the degradation rate of these gaskets.

All previous performance qualifications and the resultant service life prediction tools were performed under hot air exposure. In an effort to qualify the accuracy of the qualification and service life prediction tools in non-oxidizing atmospheres, PVRC project 99-04G was commissioned. This project investigates the effect of long term pressurized hot nitrogen exposure on the performance of foil reinforced flexible graphite type gaskets and corrugated insert flexible graphite gaskets. Comparisons were made with the long term pressurized hot air test results of earlier PVRC projects and a method was established that can adjust the PVRC qualification tools for the calculation of long term gasket temperature and service life for these materials when employed in non-oxidizing environments. The method is based on the establishment of a recommended Ag (Equivalent Aging Parameter for Flexible Graphite) value that is specific for an internal fluid exposure condition.

OPTIMIZED/HARMONIZED ROOM TEMPERATURE LEAKAGE TEST PROCEDURE FOR GENERATION OF ASME & CEN CODE GASKET FACTORS – ANALYTICAL PHASE
Warren Brown, Luc Marchand, Michel Derenne

WRC Bulletin 508
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January 2006 ISBN: 1-58145- 515-1 76 pages
This bulletin reports a common core room temperature leakage test procedure for generation of design data for both the draft ASME Non-mandatory Appendix BFJ and CEN EN 1591-1. The work has been prepared from the results of over 400 ROTT tests performed since 1994 according to the proposed ASTM ROTT Draft 9 [7] or Draft 10 [3] methods. There was no CEN prEN13555 test data available to perform an analysis of the method, as was done for the ASTM ROTT Draft methods in this report. The differences between the draft ASTM and CEN prEN-13555 methods (points of harmonization) are outlined following. There are, in reality, only three differences between the requirements placed on the ROTT gasket test procedure by the draft ASTM and CEN prEN-13555 methods. These differences listed below are: 1) fluid pressure level, 2) gasket stress levels and test sequence, 3) gasket size. For each difference, a conclusion from this report regarding the proposed method of harmonization is included. In addition, a measure of the level of variation or error associated with the proposed harmonization change is also listed. The level of variation associated with that aspect of harmonization should be compared with overall levels found in the ASTM ROTT / ASME-BFJ


ANALYSIS OF THE EFFECTS OF TEMPERATURE ON BOLTED JOINTS
Warren Brown
WRC Bulletin 510
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WRC Bulletin 510 –April 2006

Thermal events and transient thermal effects are known to play a major role in pressurized flanged joint leakage. When a leak occurs the engineering challenge is to understand and properly diagnose the role temperature and thermal transients in that failure and thereby specify measures necessary avoid future leaks. Also, since most pressure vessel codes require the consideration of thermal effects without providing the methodology, perhaps the greater engineering challenge is to the flanged joint designer. This is not only to determine temperature effects on flanged joint designs to comply with code requirements but also an evaluation of the design to assure leak free operation considering anticipated thermal events. This Bulletin provides a set of analytical tools and guidelines for addressing these challenges. The purpose of PVRC Project 01-BFC-05 was to summarize the findings of the Author’s doctoral project on the effects of temperature loads on bolted flange joints and publish a summary document providing design guidelines to deal with temperature effects and the relative magnitude of these effects. Reporting on this work, the Bulletin provides users with simplified thermal calculation methods that enable the flanged joint designer or field troubleshooter to determine the effect of steady state or transient temperatures on flanged joints, including an evaluation of leakage possibilities.

This is accomplished by the author in a readable step by step process that provides the tools to answer along the way questions such as: What increase in assembly bolt load would be sufficient to overcome anticipated thermal events? Is flange deflection caused by joint component thermal interaction sufficient to cause a leak from loss of gasket load? Could radial shearing of the gasket from differential radial expansion of the flanges or tube-sheet caused either by differences in mating flange temperatures or material properties result in failure? Is gasket crushing a possibility due to increased load caused by joint component thermal interaction? How much bolt and gasket load could be lost because of a process thermal transient, or a sudden cool-down, without failure? If insulation is applied to an operating un-insulated flanged joint, how much hotter than the flange ring might the bolts be? Is this transient sufficient cause a leak?

Following an overview of bolted flanged joint response to thermal loads, causes of failure and a background description of mechanical and thermal analysis, a detailed calculation procedure is presented. For simplicity axisymmetric and generally identical mating flanges are assumed by the calculation method. Detailed guidance is then provided on extending this approach to non-identical flange pairs, joints with a tube-sheet (Heat exchanger girth joints) and coverplates. The method first provides steady state thermal and deformation results followed by means to calculate the steady state bolt load using component compliance. An evaluation of transient effects by extension of these findings and the use of a series of graphs providing the time to reach 95% and 5% the steady state temperature for each component completes the process. Appendices A through D provide additional information on the details of the calculation methodology. Appendices E and F illustrate the calculation method via an available Excel spreadsheet.

 

[gvc99]

As I recall, plasticity analysis of annular bearing around a circular hole indicates that the bearing stress can be significantly higher than the yield stress before plastic flow occurs. Perhaps you can use this fact to allay the loss of preload concern.