Hibernia Study (Flare)

HIBERNIA MANAGEMENT AND DEVELOPMENT COMPANY LTD

FLARE SYSTEM REVALIDATION STUDY

TECHNICAL NOTE

DOCUMENT NO : 8266-HIB-TN-C-0001

REVISION : B

DATE : October 2000

8266-HIB-TN-C-0001 Page of 139 Revision: B/tt/file_convert/55276c7749795994178b46d5/document.doc October 2000

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DOCUMENT REVISION RECORD

REV DATE DESCRIPTION PREPARED CHECKED APPROVED

Draft 21/07/00 Issued for IDC M. Goodman A. Robinson M. Goodman

A 23/08/00 Issued for Comment / Final M. Goodman A. Robinson M. Goodman

B 16/10/00 Final M. Goodman A. Robinson M. Goodman

RELIANCE NOTICE

This report is issued pursuant to an Agreement between Granherne (Holdings) Limited and/or its subsidiary or affiliate companies (“Granherne”) and HIBERNIA MANAGEMENT AND DEVELOPMENT COMPANY LTD which agreement sets forth the entire rights, obligations and liabilities of those parties with respect to the content and use of the report.

Reliance by any other party on the contents of the report shall be at its own risk. Granherne makes no warranty or representation, expressed or implied, to any other party with respect to the accuracy, completeness, or usefulness of the information contained in this report and assumes no liabilities with respect to any other party’s use of or damages resulting from such use of any information, conclusions or recommendations disclosed in this report.

Title:

FLARE SYSTEM REVALIDATION STUDY

QA Verified:

TECHNICAL NOTE Date:


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CONTENTS

FRONT PAGE

DOCUMENT REVISION RECORD

CONTENTS

ABBREVIATIONS

HOLDS

1.0 INTRODUCTION

2.0 SUMMARY AND CONCLUSIONS

2.1 Introduction

2.2 Technical Audit of the Design Calculations

2.3 Challenge Process

2.4 As-Building the Flare System

2.5 Risk Management in Relation to Wind Condition and Flaring Events

2.6 Implications for Hibernia

2.6.1 Introduction

2.6.2 Capacity Opportunities

2.6.3 Impact on the Design Documentation

2.6.4 Future Work

3.0 DESIGN BASIS

3.1 Introduction

3.2 Safety Design Basis

3.2.1 Probabilistic Design Criteria

3.2.2 Deterministic Design Criteria

4.0 APPROACH

4.1 General

4.1.1 Flare System Revalidation Process

4.1.2 Legislative Obligations of HMDC’s Safety Design Philosophy

4.1.3 The Requirements of the Standards, Codes of Practice and Recommended

Practices

4.1.4 Ambiguities in the Recommended Practices

4.2 Calculation Audit


4.4 Risk Management in Relation to Flaring Events and Wind Condition


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5.0 TECHNICAL AUDIT OF THE DESIGN CALCULATIONS

5.1 Introduction

5.2 Results of the Technical Audit – Relief and Blowdown System Calculations

5.2.1 Technical Audit Issue Discussion – Relief and Blowdown System Calculations

5.3 Results of the Technical Audit – Relief Valve Sizing Calculations

5.3.1 Technical Audit Issue Discussion – Relief Valve Calculations

5.4 Technical Audit Conclusion Summary

6.0 CHALLENGE PROCESS

6.1 Introduction

6.1.1 The Principles of the Legislation

6.1.2 Relevant Canadian Legislation

6.1.3 Applying the Legislation

6.2 Jet Fire

6.2.1 Requirements of the relevant regulations, design codes and practices when

Hibernia was designed

6.2.2 How Jet Fire Was Actually Handled During Design

6.2.3 Current Requirements of the Design Codes and Practices

6.2.4 Current Best Industry Practice

6.2.5 The Effect of Applying Current Best Industry Practise to Hibernia

6.2.6 Jet Fire Conclusions

6.3 Blowdown (Depressuring) System Sizing

6.3.1 Requirements of the Codes, Guides, Standards and Recommended Practices

When Hibernia was Designed

6.3.2 How the System was Designed

6.3.3 Current Requirements of the Codes and Recommended Practices


6.3.5 The Effect of Applying Current Industry Practice to Hibernia

6.4 Compressor Blowdown Stagger







6.5 Two-Phase Relief



6.5.2 How the System was Designed for Two-Phase Relief

6.5.3 Current Requirements Of The Codes And Recommended Practices

6.5.4 Current Best Practice


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6.5.5 The Effect of Applying Best Industry Practice to Hibernia

6.6 Design Windspeed and Direction


when Hibernia was Designed

6.6.2 The Windspeed Used During Design




6.7 Acceptable Flare Radiation Levels

6.7.2 The Radiation Levels Used in the Design




6.8 Challenge Issues Resulting from the Technical Audit of the Design Calculations

6.9 Miscellaneous Issues

6.9.1 Insulation

7.0 AS-BUILDING THE FLARE SYSTEM

7.1 Introduction

7.2 As-built and Design Capacity

7.2.1 Case 1 - Design Blowdown Rate (as RABS Rev C1)

7.2.2 Case 2 - ‘As-Built’ - i.e. As Case 1 with Future Equipment Removed

7.2.3 Case 3 - As Case 2 with 3 min Stagger Removed

8.0 RISK MANAGEMENT IN RELATION TO FLARING EVENTS AND WIND CONDITION

8.1 Introduction

8.2 Potential Flare Envelope based on Total Blowdown Scenarios

8.2.1 Determination of Blowdown Load Basis

8.2.2 Determination of Continuous Load Basis

8.2.3 Determination of Allowable Thermal Radiation Impingement on the Platform

8.2.4 Other Calculation Criteria

8.3 Results

8.3.1 Emergency Relief - Platform Blowdown

8.3.2 Continuous Relief

8.4 Flare Envelope Conclusions

8.4.1 General




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9.0 IMPLICATIONS FOR HIBERNIA

9.1 Introduction

9.2 Flare System Capacity Opportunities

9.3 Impact on the Design Documentation

9.4 Optional Changes

9.5 Miscellaneous Requirements

9.5.1 Updating the Design Documentation

9.5.2 Implementation Projects

10.0 REFERENCES

APPENDIX I CALCULATION TECHNICAL AUDIT

APPENDIX II STAGE 2 PROPOSAL


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ABBREVIATIONS

ALARP As low as reasonably practical

API American Petroleum Institute

ASME American Society of Mechanical Engineers

CSE Concept Safety Evaluation

CBA Cost Benefit Analysis

DAE Design Accidental Event

DIERS Design Institute for Emergency Relief Systems

DPRA Design Phase Risk Assessment of Potential Accidental Events

DPSEE Design Phase Safety and Environmental Evaluation

FRA Fire Risk Assessment

FSRS Flare System Revalidation Study (i.e. this study)

H&S Health and Safety

HEM Homogeneous Equilibrium Model

HSE Health and Safety Executive (UK government body)

HSW Health and Safety at Work

HTPT Hibernia Topsides Process Team

HVAC Heating Ventilation and Air Conditioning

MEP Mobil Engineering Practice

N North

NW North West

PFP Passive Fire Protection

PSHH Pressure Switch High High (a trip function)

QRA Quantified Risk Analysis

RABS Relief and Blowdown Study Report Doc. No. CM-E-C-R-M00-RP-3410 Rev C1

RAE Residual Accidental Event

TSR Temporary Safe Refuge

VLE Vapour Liquid Equilibrium


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HOLDS

1. No holds.


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1.0 INTRODUCTION

The key Hibernia flare system design documents have remained unaltered since the

design phase. In certain areas, the key assumptions are now considered worthy of

review particularly to incorporate as-built system details and to assess the potential to

remove the Injection Compressor stagger. It has therefore been decided to revalidate

the key flare system design documents (principally the Relief and Blowdown Study

Report and the Flare System Calculation Volumes). Following from a series of

meetings during the period 9–10 May 2000, a scope of work for performing a staged

revalidation of the flare system was prepared.

The stages envisaged are described below:

Stage 1 – Flare System Revalidation Report

Stage 1 consists of the following activities:

Review the original documents in light of best practice to ensure a consistent and

clear design approach (including a technical audit of the existing flare system

design calculations). Analyse the results of any changes in design philosophy and

their impact on flare system design capacity.

Based on the above identify the changes required to update the Relief and

Blowdown Study Report.

Identify the available capacity in the system against a range of future projects

including the avoidance of Injection Compressor blowdown stagger.

Prepare a discussion document describing the background and requirement for any

changes to the key flare system documents.

Stage 2 – Modify the Key Hibernia Flare Design Documentation

Based on the results of Stage 1 update the key flare system design documentation.

This report covers Stage 1 of the revalidation process.


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2.0 SUMMARY AND CONCLUSIONS

2.1 Introduction

The key design documents relating to the flare system design have not been reviewed

for some time. In the intervening period codes of practice have changed, as-built

documentation and better analytical tools have become available and future

equipment, foreseen at the time of design, is no longer certain to be installed. This

report addresses these issues in order to revalidate the flare system design.

The results of the report are summarised in the following main sections:

Technical audit of the design calculations

Challenge process

As-building the flare system

Risk management in relation to wind condition and flaring events

Implications for Hibernia

These are described in turn below.

2.2 Technical Audit of the Design Calculations

The technical audit of the design calculations was undertaken primarily to identify

assumptions which were linked to the relief and blowdown system design basis and

because of some concern in HMDC that the calculations did not fully reflect the

design. In the event very few important assumptions were contained in design

calculations but it was clear that the calculations were not up to date and there were

some inconsistencies between the various flare system design documents. Also

design methods have improved which indicate certain assumptions are no longer

sufficiently conservative (for instance in the low temperature material selection

calculations).

Otherwise, as would be expected, some of the design bases on which the system was

founded have changed since the design phase and an exercise such as this is the

ideal way of capturing these changes (for instance the changes relating to the

maximum well rate).

One last aspect uncovered in the technical audit related to missing work (for instance

the LP flare network model which should have been run to calculate the back pressure

on relief valves which may be part of a coincident relief).

All these types of issue will require the design calculations to be revised and corrected.

A summary of the calculations that require revision can be found in Section 2.6.


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The challenge process showed the problem of applying, often ambiguous, codes of

practice retrospectively. The process also showed that Hibernia has in general used a

conservative design basis which has resulted in a robust design when considering new

code requirements and current industry design approaches. Generally, this is

because newer approaches tend to interpret the codes without incorporating

unnecessary features, which reflects efforts to achieve low cost facilities. Only in one

area did we believe the design had not been sufficiently conservative and this area

was the use of compressor stagger to limit the LP flare system flowrate during

blowdown. However, to remove the stagger could present a considerable design

challenge because of the problem of increasing the pressure (and inventory) in the LP

separator when the blowdown valves opened. This is particularly undesirable if the

cause of the blowdown is a fire around the LP separator. The higher pressure and,

therefore, higher stress will increase the risk of premature failure. Consequently a

brief safety analysis which looked at the ability of the A train injection compressor

components (the equipment whose blowdown is delayed) to survive an adverse fire

was undertaken. The results demonstrated the equipment is unlikely to fail and,

therefore, the staggered system is safe in this situation.

Similarly, the challenge process also considered jet fire on lower pressure equipment.

Here it was found that thin walled vessels should be considered outside of the API

guidance (as suggested explicitly in the API recommended practices). In this case

applying modern practices relating to jet fire to this vessel suggests that as long as the

insulation remained intact the insulation will ensure the vessel survives a jet fire.

One other area, which would have been unknown to the original designers, are the

changes which have occurred to the recommended practices; here the only important

change relates to new API sizing method for calculating two-phase relief which will

need to be applied to the existing system.

The outcome of the challenge process is described in Section 2.6.

2.4 As-Building the Flare System

The removal of the allowances contained in the design for future equipment ‘frees-up’

approximately 30% of the HP flare capacity in the blowdown case. However there is

no impact at all on the LP flare capacity for this case as no future equipment was

planned to be connected to this system.


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2.5 Risk Management in Relation to Wind Condition and Flaring Events

We had expected to find some useful work regarding this aspect in the industry but

found none. Generally a picture emerged that the industry has no standard approach

to windspeed or the operational measures undertaken on a platform during high wind.

The outcome of such considerations is therefore undertaken on a case by case basis

relying mainly on the judgement of the asset owner. By applying a consistent

approach to the selection of design windspeed and the presence of personnel the

result changed the design case relatively little. The impact of the design windspeed

change, if pursued, is summarised in Section 2.6.

One aspect where consideration of wind condition would potentially have a beneficial

effect is related to continuous case flaring. In this case, because of the lower

allowable radiation levels, wind has a significant effect on what can be produced when

the compressors are unavailable. This suggests an allowable flare radiation envelope

should be developed such that the production rate can be set (maximised) dependent

on the measured (or expected) wind speed and direction.

2.6 Implications for Hibernia

2.6.1 Introduction

The implications for Hibernia effectively arise as two types.

Where analysis suggests that some aspects of the design are conservative

compared to the application of recommended and best practices, then this implies

some apparent spare capacity in the system. The use of such capacity is optional

dependent on future plans for the facility and economic benefit.

However, where analysis suggests that some aspects of the design are less

conservative than in the recommended or best practices, then this implies that the

system capacity is insufficient or marginal in these cases. In this case prudent

ownership requires that these issues are addressed and solutions provided.

2.6.2 Capacity Opportunities

The flare system revalidation analysis has identified a number of areas where a

capacity opportunity is available. Where the capacity opportunity is positive (i.e. the

apparent capacity, or capacity available in the flare system appears to rise) then

HMDC have the choice of adopting the new philosophy which can be used to allow

future platform modifications.

However, where the capacity opportunity is negative (i.e. the apparent capacity, or

capacity available in the flare system appears to fall) then HMDC should undertake

remedial measures to remove the possibility of exceeding the platform flare system

capacity.


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Both types of capacity opportunity are described below.

Changes that reduce the apparent capacity of the flare system

The new API method for calculating two-phase relief should be applied to the

relevant cases. At the same time, a new maximum design well rate together with

the number of wells which fail to shut in that need to be designed for will need to

become an explicit part of the RABS update. If the existing relief valves are to be

retained this may require some method for limiting the maximum well rate to an

acceptable value.

The missed design case of a failed open separation train spillover valve should be

calculated and measures sought to limit the peak rate experienced to within the

capacity of the flare system.

Changes which apparently increase the capacity of the flare systems

Remove the effect of future equipment.

Change the start pressure of the commencement of blowdown. A robust

interpretation of the codes of practice suggests in an automatically initiated

blowdown event the start point should be normal operating pressure. Some of the

Hibernia systems already follow this philosophy, however some (the compressor

systems) begin blowdown from the PSHH pressure (which is significantly in excess

of normal operating pressure). Changing this to be more consistent would free-up

capacity in the LP flare system. However, in doing so there is the disadvantage of

having to use more rigour when considering any changes in normal operating

pressures, which affect the blowdown and which the PSHH approach can avoid.

Increase the end pressure for blowdown for the thick walled vessels. The API

recommended practice has never required blowdown to 690 kPag for thick walled

vessels. Our calculations of vessel heat up confirm these conclusions. Therefore

blowdown to 690 kPag is excessive in this case. The end pressure should be 50%

of the design pressure unless there are good reasons otherwise. One such

situation is the inapplicability of this end pressure for the gas turbine driven HP gas

compressors. The blowdown end pressure for equipment in these systems is

determined by the requirements of the HP compressors seal oil system and the

requirement to be at or near atmospheric prior to the oil running out.

The effects of insulation. If the insulation is the right type and properly attached its

effects can properly be considered in the design calculations.


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Radiation levels. The Hibernia radiation levels did not truly reflect how the platform

was constructed. In particular the radiation requirement on the drilling derrick does

not reflect the shielding present, nor do the radiation levels on the escape ways

reflect international practice or Canadian regulations. The allowable radiation in

these levels should be raised to 9.5 kW/m2 and 6.3 kW/m2 respectively. This

allows considerably higher flaring rates before the radiation levels are breached.

Windspeed. Using a probabilistic method to select windspeed would lower the

design windspeed from 27 m/s to 24.2 m/s.

Gas blowby cases are over conservative. By taking a more realistic blowby case

the apparent capacity can be increased. This would pay benefits should the

separation train LCV valve coefficients (Cv) ever need to be raised.

The detailed analysis which forms the basis for the above can be found in sections

5.0, 6.0 and 9.0.

2.6.3 Impact on the Design Documentation

The following table summarises the changes that should be made to the design

calculations to improve their integrity and make them consistent and traceable.


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Table 2.1 Calculations Requiring Revision (System 34)

Number

34-

Title Number

34-

Description Action

005 / A Blowdown Section Inventory Calc (Provides input to blowdown simulations)

005/1 Are the blowdown volumes used sufficiently accurate?

Locate and review missing

calculations

005/3 Were the real settle out pressures ever

used?

Compare real settleout conditions

with design to ensure blowdown

rates are appropriate

006 / A Blowdown Summary 006/1 HP Blowdown calculation higher than

vendor aware of. Radiation level for

case is underestimated.

Update RABS.

006/2 Correct isentropic efficiency used? An optimistic isentropic efficiency

was used to calculate the

minimum system temperature.

Recalculate the temperatures.

See also 34.010/1.

010 / A Calculation of allowed

cooldown before

hydrate formation &

minimum

temperatures

achieved in flare gas

from critical blowdown

sections

010/1 Was the calculation methodology

sufficiently robust?

There are flaws in the method

used to calculate the minimum

temperatures in the system.

These should be corrected. Use

resultant more realistic figure to

implement alarms on high

pressure areas to avoid low

temperatures. Update RABS.

011 / A Review of HP flare

KO Drum size

011/1 A note on the front of calc 34-064 states

that Rev 7 of Design Basis gives max

well flow of 20,000 bpd + average well of

10,000 bpd, i.e. 30,000 bpd total. The

individual well design rate has changed.

What are the implications for the

platform?

Select number and design rate of

the well failure to shut in case.

Update RABS. Develop

operational procedure to cater for

time to fill HP flare KO vessel.

015 / A Calc to review options

for reducing HP to MP

Separator and MP to

LP Separator Blowby

Cases

015/1 Relief & Blowdown Study Report Rev C1

non-concurrent maximum allowable LP

and HP Flare loads are 110,874 kg/h

and 244,897 kg/h respectively. Rates

used in these calculations exceed

design.

Ensure design rates quoted are

consistent and reflect the installed

control valves. Update RABS.

022 / C HP Flare Network

Sizing (HP Separator

- Max Relief Case)

022/2 Effect of increased production /

production fluid GOR

Update RABS to mention link

between GOR and the compressor

capacity.

025 / C 3rd Stage

Compressor Max

Relief Case - Network

Analysis

025/2 Include in updated RABS cases which

are not catered for, i.e. consider relief

from both compressor trains

Check modifications to avoid

injection compressor RVs lifting

prevent coincident case. Update

RABS to explicitly mention the

cases which are not designed for.

033 / G Coalescer & LP

Separator Heaters

Simultaneous Fire

Relief - Network

Analysis

033/1 Assumption that the header is at zero

pressure (I.e. that this is a singular event

not coincident with any other releases)

Construct an LP flare network

model to calculate the back

pressure on relief valves when the

system is depressuring.


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Number Title Number Description Action

31.37 Relief Valve

Calculations - LP

Separator

31.37/1 Is it possible for the Test Separator

manifold to be connected to the LP

Separator when operating in high

pressure mode?

Ensure positive method of

ensuring isolation from HP system

exists. Update RABS to reflect

this.

The tables avoid repetition of the major issues which affect the capacity of the flare

system (see Section 2.6.2) and single issues that affect more than one item. The full

version of the tables can be found in sections 5.2.1 and 5.3.1.

Some minor changes which generally affect consistency are identified in Section 9.4.


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2.6.4 Future Work

It is difficult to be precise regarding the activities required in the longer term future as

these depend on the outcome of the recalculation work and the decisions made

therein. However, it is possible to summarise in general terms the shorter term

requirements of Stage 2:

Document Changes

Relief and Blowdown Study Report

Fairly extensive rewrite of the report.

Design Calculations

For each change prepare a calculation revision which revs up the existing

calculation (in other words building on the existing work). This would include:

- Calculations identified in this report requiring change.

- Flare radiation calculations (for windspeed and allowable radiation levels)

- Continuous radiation cases. Analysis of allowable production rate vs wind

speed and direction.

Blowdown inventory calculations (for removed inventory).

Reliability analysis of the system that controls the compressor stagger, to ensure

the system is sufficiently reliable to ensure the design integrity.

Implementation Projects

In this section there are some projects mentioned which will in all likelihood require

hardware changes to be made (resulting from the above there may be more).

Insulation conformance - The explicit ability of the platform to cope with a jet fire

hazard requires the insulation around the vessels to remain in place during jet

flame impingement. This may require the insulation strength to be improved.

Modifications to limit peak flaring rate during spillover valve failure.

Instrument modifications to warn operators when the requirement to blowdown

compressors is becoming imminent (to avoid low temperature problems).

In discussion with HMDC a detailed scope of work to undertake the above has been

developed. This is attached in Appendix 2.


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3.0 DESIGN BASIS

3.1 Introduction

The review focuses on those aspects and hazards that are directly relevant to the flare

system design. In particular these are flare radiation, jet and pool fire and, to a lesser

extent, explosion. This naturally excludes issues relating to blowout and iceberg

collisions as well as environmental issues.

However, before going into the main parts of the analysis, it is worth recapping the

safety basis and methodology followed by HMDC now and during the design phase.

This will need to be followed should any changes be made to design philosophies.

3.2 Safety Design Basis

As was convention at the time, the safety design progressed along two parallel routes.

The first was the use of probabilistic analysis to identify the acceptability of various

risks. The second was the deterministic design of the various safety systems

according to recognised codes and practices. Occasionally there was an interface

between the two processes when a risk was considered unacceptable. Where this

was apparent the design would be adjusted to mitigate the unacceptable risk.

These two processes are described in sections 3.2.1 and 3.2.2 below.

The problem of parallel processes is that some information can be lost across the

interface. More recently this has led to a concept called risk based design where the

key safety issues are resolved during the early conceptual design stages rather than

be left for implementation after the conceptual design is complete.

3.2.1 Probabilistic Design Criteria

By the time the FRA was commenced the HMDC Damage / Impairment Criteria had

been formalised. These were:

Criterion 1: Overall Platform Integrity

There must be no overall loss of integrity of the platform for at least 2 hours after the

initial event. Loss of integrity included:

Structural collapse of more than 50% of the platform topsides, or total collapse of

Module M30.

The 2 hours is judgementally used for a maximum time to evacuate by lifeboats (i.e.

time to respond to emergency, attempt to control, organise evacuation and abandon

platform)


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Criterion 2: Integrity of Temporary Safe Refuge (TSR)

The TSR (i.e. living quarters) should remain a safe refuge for personnel for at least 2

hours. Loss of integrity may be due to:

Fire within the safe refuge;

Blast damage, in excess of major window breakage;

Collapse of any part of shelter area.

The time of 2 hours is derived and defined as for Criterion 1.

Criterion 3: Escape Routes

Escape routes from all parts of the platform to the TSR or other safe refuge should

remain passable for at least 30 minutes from the start of the incident. An escape route

may be made impassable by:

Thermal radiation over 12.5 kW/m2 to the outside of the escape route if protected by cladding;

Thermal radiation over 6.3 kW/m2 if unprotected;

Blockage due to blast damage;

Collapse of one or more modules;

Flooding over 1 m deep in the Utility Shaft.

The time of 30 minutes is intended to allow the escape of workers who had initially remained at their posts to shut-down the process operation or fight a growing fire. The criterion is violated if an incident results in either:

All escape routes from any module being impassable;

All routes from any one part of the platform to the TSR being impassable.

Since there is more than one escape route from any point, an incident must completely involve the total module to violate the criterion.


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Criterion 4: Means of Evacuation

The evacuation systems must remain effective for long enough to evacuate all

personnel. This requires at least one of the following to be true:

Helideck operable for at least 2 hours. Inoperability may result from one of the

following:

tilt over 15°

smoke due to oil fire and wind towards helideck

thermal radiation over 3.2 kW/m2

blast damage

unignited gas over helideck (due to likelihood of ignition)

collapse of M50 module.

Evacuation systems must be operable with at least 10% spare capacity (to allow for

launching partly loaded) and with passable escape routes, from TSR to evacuation

system, for at least 2 hours. Inoperability may result from:

tilt over 25° (preventing safe access)

thermal radiation over 12.5 kW/m2

blast damage (damaging launching gear and access walkways)

collapse of module M40 and M30.

The times are judgementally based on the proposed systems for the Hibernia platform.

The criterion is only violated if all the means of evacuation are unavailable.

In general the following approach was applied:

The Damage / Impairment Criteria set out above, give basic criteria which should not

be exceeded. It is not possible to ensure that no incidents will exceed the criteria.

The intent is that every reasonable and practical precaution is taken to ensure those

incidents that exceed the criteria are so unlikely that they can be considered as an

acceptable risk because the risk is negligible. These incidents are termed Residual

Accidental Events.


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Added to this were Hibernia’s three-tier framework of risk acceptability:

For any single incident that might affect the key safety systems (more accurately

safety functions from the above), the risk level for the three-tiers are:

Intolerable: greater than 10-4 per year.

ALARP region: 10-4 to 10-5 per year.

Lower bound of acceptability: less than 10-5 per year

Whilst it is inconceivable that any of the impairment criteria would change as a result

of the considerations in this report, change may affect the QRA upon which these

impairment criteria stand. Any changes considered, therefore, will require to be

confirmed through QRA.

3.2.2 Deterministic Design Criteria

A number of the final requirements for the design would stem from the above. This is

not surprising as some of the aspects of the impairment criteria actually have their

roots in the recognised international codes and practices, e.g.

Allowable Flare Radiation Levels:

Escape Ways

Not over 12.5 kW/m2 to the outside of the escape route if protected by cladding;

Not over 6.3 kW/m2 if escape way is unprotected

Helideck

Not over 3.2 kW/m2

The remaining requirements are part of the various design guides and codes of

practice. These are described in detail in Section 6.0


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4.0 APPROACH

4.1 General

4.1.1 Flare System Revalidation Process

The flare revalidation process is summarised in the flowscheme overleaf. The stage

consisting of this report is Stage 1. The results of Stage 1 will form the basis for future

Stage 2 studies. The flowscheme indicates the potential range of projects this could

encompass.


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Figure 4.1 - Revalidation Process

Start

Technical audit of the existing

design calculations Challenge process Risk management

Prepare Flare System Revalidation

Report

Draft and Final Versions

Impact on

the design?

Update Relief and

Blowdown Study Report

(Rename “Flare System

Design Philosophy”)

Build flare network model

(for inclusion in the Flare

System Design

Philosophy) (Optional)

Stage 1

Major changes to flare

design calculations. As

build and replace the

design calculation volumes

Update Relief and






System Design Philosophy)

Prepare workscopes for the

modifications

Minor changes to flare

design calculations. Rev up

affected calculation

volumes

Update Relief and






System Design Philosophy)

(Optional)

End

Major impact Minor impact

Stage 2

Varies according

to outcome of

Stage 1

No Impact

HMDC Internal Audit /

Approval

HMDC Internal Audit /

Approval


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4.1.2 Legislative Obligations of HMDC’s Safety Design Philosophy

HMDC’s legislative obligations to safety are encompassed in the Newfoundland

Offshore Petroleum Installations Regulations (Reference 1), an extract of which

follows:

43. (1) Every operator shall…submit to the Chief a concept safety analysis…that

considers all components and all activities associated with each phase in the life of the

production installation, including the construction, installation, operation and removal

phases…

(5)…

…(g) a definition of the situations and conditions and of the changes that would

necessitate an update of the concept safety analysis.

(8) The operator shall maintain and update the concept safety analysis referred to in

subsection (1) in accordance with the definition of situations, conditions and changes

referred to in paragraph (5)(g) to reflect operational experience, changes in activity or

advances in technology.

HMDC have met these requirements, initially through the preparation of the Concept

Safety Evaluation which has, over time, evolved into the Operational Plan which will be

issued in the near future.

The above very much parallels the type of approach the UK HSE require:

11. The employer…needs to review the risk assessment if there are developments

that suggest that it may no longer be valid (or that it can be improved). In most cases,

it is prudent to plan to review the risk assessments at regular intervals - the time

between reviews being dependent on the nature of the risks and the degree of change

likely in the work activity. Such reviews should form part of standard management

practice.

Management of health and safety at work - Approved Code of Practice (1992)

The study will be undertaken in this light.


24

4.1.3 The Requirements of the Standards, Codes of Practice and Recommended

Practices

Where a particular design code is used its requirements are mandatory, e.g. the

requirements of ASME VIII for a vessel stamped accordingly. Recommended

practices are different in that their requirements are not mandatory in law unless they

are stated in the regulations; this is the case with Canadian requirements. For those

practices not in the regulations, common industry practice and deviation from those

would normally require Certifying Authority approval. In dispute, the applicability of the

use of the practice would be left to the courts to decide.

So, although recommended practices are not the same as design codes, their

requirements have become almost code like over time. Consider a situation where a

failure has occurred and its cause appears to be linked to a situation where a well

known, recommended practice was deviated from. There would in effect be an onus

to prove the requirement inferred in the practice was inappropriate at the time it was

being considered. Otherwise negligence would be very difficult to disprove. This proof

would be particularly hard to provide and would require thorough documentation

regarding the deviation to be kept for the life of the plant. This study will recognise this

reality and, therefore, design code, code of practice and recommended practices, so

long as they emanate from a recognised responsible body, are considered equivalent

in this study in terms of reliance.

4.1.4 Ambiguities in the Recommended Practices

The recommended practices are sometimes (some would say often) ambiguous in

their requirements and a study such as this tends demonstrate the problem. Therefore

interpretations of the practice’s actual intent often have to be made. This is one of the

designer’s challenges and a particular challenge of this report.

In the past, presumably to avoid the need for interpretation, owners of facilities have

removed the ambiguities by being prescriptive with their requirements. Thus, industry

practices, which often have little connection with the original design code intent, have

sometimes been established for expediency. The goal today is to apply the codes as

intended without unnecessary features that increase cost.

The methodology used in this study therefore lies in this latter approach to the

application of practices, to apply them as intended. Where there are differences

between the various methods of applying practices this will be highlighted in the

narrative.


25

4.2 Calculation Audit

The calculation technical audit’s primary aim is to identify key assumptions or queries

contained outside the Relief and Blowdown Study Report. A secondary aspect is the

high level numerical and methodological check of the existing relief case calculations

to ensure their suitability to use as the basis for the revalidated Flare Design

Philosophy. This process will identify any areas that require detailed review to be

undertaken in future stages of the revalidation study.

The audit will use a tabular approach (compiled by calculation volume) to highlight the

assumptions or issues which require to be addressed during the challenge and risk

mitigation review processes in subsequent sections of the study.

Whilst addressing the calculations, secondary objectives such as consistency and

methodology have also been revisited and these results can also be found on the

detailed audit sheets.


The challenge process is a sequential review of individual design parameters that had

an effect on the way the flare system was dimensioned.

The process looks first at the requirements of the design codes or practices to

place, in an historical context, the requirements for the design. The key codes and

practices applicable to this work are those referred to in the RABS, i.e.:

Canada Oil and Gas Installation Regulations

Petroleum Occupational Safety and Health Regulations - Offshore

Newfoundland

API RP 520

API RP 521

API RP 14C

Mobil EGS 661

ASME Section VIII, Division 1

Secondly, how the system was actually designed is considered. This aspect

captures the interpretations used when compared to the earlier activity.

Thirdly, the requirements of current codes and practices are then considered.


26

Fourthly, current best practice is considered to challenge the existing design and

determine its suitability for application to Hibernia. Here the best practice applied

is Granherne’s own (there is no other convincing way for us to address this issue).

This is not to say that other companies do not apply the requirements differently.

Where possible some alternative applications will be mentioned where relevant.

Finally, the effect of these stages is considered and a recommendation made for

the way the requirements should be applied to Hibernia to give a consistent and

easily understood flare system design.

4.4 Risk Management in Relation to Flaring Events and Wind Condition

This issue stems from the technique used occasionally where the capacity of a flare

system has been increased when it has been realised that at the design windspeed no

personnel would be present on deck, thereby allowing higher incident radiation rates

on deck during these events. This would only be the case if a very high wind speed

were considered during design. A more pragmatic approach to design windspeed

selection would ensure that the coincident conditions were considered, i.e. a

realistically high windspeed.

Generally two cases are considered:

Emergency flaring

The methodology used to generate the results of the activity are based on multipliers

applied to the flowrates considered in the original flare boom length defining design

case, i.e. combined HP and LP blowdown. It should be appreciated that these

resultant rates cannot actually occur until the platform is actually modified to

incorporate the necessary inventory, in this case, in the same ratio (HP/LP) as design.

This is unlikely. The real effects, or real envelope, will therefore depend on the actual

modification made. The actual effect should be considered in detail during the

particular modification project’s design phase.

Continuous flaring

Continuous flaring differs from emergency flaring in that the acceptable radiation levels

are very much lower than for emergency events. Proportionally, the lower radiation

isopleths are more sensitive to wind than the high flow cases.

The methodology used in this study has taken a simulation from the recent

debottlenecking project (Case 3, a case which included Avalon production) and used

this to generate current normal operating input data to generate a new set of profiles.

The input data is adjusted up or down using simple multipliers to generate the

expected envelope.


27

5.0 TECHNICAL AUDIT OF THE DESIGN CALCULATIONS

5.1 Introduction

The intent of this review is twofold:

To identify assumptions or items which have affected the design of the relief and

blowdown system including any issues which arise as a result of the review

(Section 5.2).

To revisit the relief valve calculations to perform a methodological and numerical

check to reconfirm the validity of the dimensioning design cases (Section 5.3).

5.2 Results of the Technical Audit – Relief and Blowdown System Calculations

In Appendix I the full results of the audit are given. The detailed tables that follow

identify a number of issues to be dealt with, either in Section 6.0, because they can be

challenged, or in this section if they are issues which concern the accuracy or

soundness of the design conclusions. For completeness, however, both sets of issues

are summarised below.

Table 5.3 Challenge Issues - Relief and Blowdown System Calculations

(System 34)

Calculation Issue

Number

34-

Title Rev Date Number

34-

Description


06 18-May-93 005/2 Jet fire scenario not taken into account for

the design of the blowdown system

005/4 Were fire areas used for total blowdown

rate?

006 / A Blowdown Summary 05 19-May-93 006/2 Correct isentropic efficiency used?

006/3 Is design case too extreme?

042 / F Total LP Blowdown - Initial Conditions

- Network Analysis

02 18-Mar-93 042/2 Validity of staggering blowdown. Were the

systems sufficiently independent?

043 / F Injection Compressor 'A' Blowdown -

Initial Conditions - Network Analysis

02 18-Mar-93 See 34-042/2

044 / G Total LP Blowdown - After 3 mins

(stagger point) - Network Analysis

01 22-Mar-93 See 34-042/2


28

Table 5.4 Technical Audit Issues - Relief and Blowdown System Calculations

(System 34)

Calculation Issue

Number

34-


34-

Description


06 18-May-93 005/1 Are the blowdown volumes used sufficiently accurate?


used?

005/5 Are vessel weights used reasonable?

006 / A Blowdown Summary 05 19-May-93 006/1 HP Blowdown calculation higher than

vendor aware of. Radiation level for case is

underestimated.

006/4 Is constant rate blowdown a valid design

method, i.e. not according to API?

006/5 'As Built' settleout pressure

010 / A Calculation of allowed cooldown

before hydrate formation & minimum

temperatures achieved in flare gas

from critical blowdown sections

02 23-Mar-92 010/1 Was the calculation methodology


010/2 Should 'troubleshooting' methanol injection

points be incorporated?

011 / A Review of HP flare KO Drum size 02 06-Feb-92 011/1 A note on the front of calc 34-064 states

that Rev 7 of Design Basis gives max well

flow of 20,000 bpd + average well of 10,000

bpd, i.e. 30,000 bpd total. The individual

well design rate has changed. What are the

implications for the platform?

012 / A Review of LP flare KO Drum size 03 10-Mar-92 See 34-011/1

015 / A Calc to review options for reducing HP

to MP Separator and MP to LP

Separator Blowby Cases

01 14-Aug-91 015/1 Relief & Blowdown Study Report Rev C1

non-concurrent maximum allowable LP and

HP Flare loads are 110,874 kg/h and

244,897 kg/h respectively. Rates used in

these calculations exceed design.

015/2 Is considering only one control valve fails

open for gas blowby case when 2 installed

in parallel realistic / allowable even with

provision of independent transmitters and

controllers?

060 / B Indicative Injection Compressor

Cooldown Calculation

01 29-May-92 See 34-010/1 and 34-010/2

061 / B Simplistic Steady State Preliminary

Review of the Annulus Rupture Relief

Flowrate

01 10-Sep-92 061/1 This case had the potential to be the

defining case for the HP flare system

(depending on installed choke valve CV)

What happened subsequently?

022 / C HP Flare Network Sizing (HP

Separator - Max Relief Case)

02 22-Mar-93 022/1 Calculated maximum pressure at PSV

discharge exceeds value on PSV datasheet

Rev C1

022/2 Effect of increased production / production

fluid GOR


29


(System 34) (Cont.)

Calculation Issue

Number

34-


34-

Description

023 / C HP Separator Max Spill-off Case -

Network Analysis

02 22-Mar-93 023/1 Calculated maximum pressure at spill-off

valve discharge exceeds value on control

valve datasheet Rev C1

023/2 Is case where valve fails fully open

considered?

See also 34-022/2

024 / C MP Separator Max Relief Case -

Network Analysis

02 24-Mar-93 024/1 Calculated maximum pressure at PSV


Rev C2

025 / C 3rd Stage Compressor Max Relief

Case - Network Analysis

01 27-Jan-93 025/1 Calculated maximum pressure at PSV


Rev C2 - Check for later revisions

025/2 Include in upated RABS cases which are

not catered for, i.e. consider relief from both

compressor trains

026 / C Injection Compressor Max Relief


01 27-Jan-93 026/1 Calculated maximum pressure at PSV


Rev C2

See also 34-025/1& 34-025/2

027 / C MP Separator Max Spill-off Case -

Network Analysis

02 24-Mar-93 027/1 Was failed open control valve considered?

See also 34-022/2

028 / D West Test Separator Max Spill-off


02 24-Mar-93 028/1 Is case where valve fails fully open

considered.

See also 34-022/2

029 / D West Test Separator Max Relief Case

- Network Analysis

02 28-Jan-93 See 34-022/2

030 / D East Test Separator Max Spill-off



considered?

See also 34-022/2

031 / D East Test Separator Max Relief Case

- Network Analysis

02 25-Mar-93 See 34-022/2

34- / E 1st Stage Compressor Spill-off Case -

Network Analysis

01 29-Jan-93 34-/1 Calculated maximum pressure at spill-off

valve discharge exceeds value on control

valve datasheet Rev C1

34-/2 Is case where valve fails fully open

considered?

045 / E Total HP Blowdown Initial Conditions

(Checks blowdown line sizes for

individual system blowdowns)

01 22-Mar-93 045/1 There is no network analysis run with

common HP Blowdown at initial conditions

045/2 Consistency error in the number and flows

in the gas injection flowlines

039 / F LP Separator Max Spill-off Case -

Network Analysis


considered?

See also 34-022/2


30


(System 34) (Cont.)

Calculation Issue

Number

34-


34-

Description

042 / F Total LP Blowdown - Initial Conditions

- Network Analysis

02 18-Mar-93 042/1 Total blowdown rate (initial rate) used in

calc less than that in Relief & Blowdown

Study Report ( 89,601 kg/h)

033 / G Coalescer & LP Separator Heaters

Simultaneous Fire Relief - Network

Analysis

01 01-Feb-93 033/1 Assumption that the header is at zero



036 / G Injection Stage Suction Scrubber PSV

- Network Analysis

01 10-Feb-93 036/1 Inconsistency on datasheet between

accumulation and 'Max Relieving Pressure'

(should be 10%)

037 / G HM & CM Expansion Drums

Simultaneous Fire Relief Case -

Network Analysis

02 01-Feb-93 037/1 Calculated back pressure (for 0152A/B)

greater than specified on datasheet - calc

considers this OK as less than 10% of set

pressure

See also 34-033/1

046 / G Fuel Gas Cooler / Heater tube rupture

relief line size check

01 02-Mar-93 046/1 ''As Built' P&IDs show bursting discs in this

service (calc considers PSVs) therefore

calc is no longer valid

050 / G 3rd Stage Suction Scrubber A (D-

3303A) PSV Discharge Line Size

Confirmation

01 02-Mar-93 050/1 Rev C2 PSV datasheet states set pressure

= 8200 kPa(g), 'As Built' P&ID shows set

pressure = 7000 kPa(g)

052 / G E-3301 Shell Side PSV Discharge

Line Size Confirmation

01 02-Mar-93 See 34-033/1

053 / G E-3303B Shell Side PSV Discharge

Line Size Confirmation

01 02-Mar-93 See 34-046/1

054 / G HP Manifold Relief - Network Analysis 01 02-Mar-93 054/1 Rev C2 PSV datasheet states set pressure

= 34,400 kPa(g), 'As Built' P&ID shows set

pressure = 34,100 kPa(g)

055 / G Simultaneous Fire Relief Case from Z-

3701 A/B, Z-3702 A/B & Z6202 A/B

(pig launchers and fuel gas package)

01 10-Feb-93 See 34-033/1

057 / G E-3701 Shell & Tube Side

Simultaneous Fire Relief Case - Line

Size Confirmation

01 12-Feb-93 057/1 Calculated back pressure exceeds that

specified on datasheet for both PSVs

See also 34-033/1

058 / G E-6201A/B Tube Side Fire Relief

Case

01 10-Feb-93 See also 34-033/1

059 / G Comparative Program check of

INPLANT Single Phase Simulation vs

ESI

01 23-Apr-93 059/1 Accuracy of calculations using ESI instead

of INPLANT


31

5.2.1 Technical Audit Issue Discussion – Relief and Blowdown System Calculations

The following describes the technical audit issues identified in Table 5.4. However,

first the general issues relating to the audit are described.

5.2.1.1 General Issues

A number of issues arose which affected many of the calculations in the flare system

calculation volumes. These are described below:

Flare network calculations - Whilst the flare network calculations were performed,

the results of the calculations were never carried over to the discipline (instrument)

data sheets (through which the equipment was purchased). In other words, the

control valves and relief valves were all sized with the wrong back pressure. In

most cases this has no effect because the relief valves are balanced and the

difference in back pressure is low or, for similar reasons, because we know the

control valves appear to be doing their respective duties (albeit probably a little

more open than planned). Where there is an effect this is noted as an issue below.

Fire zones were used in the calculations but Granherne, so far, have not had

access to documents describing them.

Vendor data didn’t make it through to the final calculations. This particularly

affected the settleout pressures for the compressors and the calculations of realistic

volumes in the system. Because of the aggregate nature of these changes we

suspect, but cannot be sure, there would be no material effect on the flare system

design.

5.2.1.2 Technical Audit Issues

Issue 34-005/1 - Are the blowdown volumes used sufficiently accurate?

The majority of the blowdown volume data is summarised and unchanged from an

earlier revision of the calculation that used the best available information at the time for

piping volumes. The separation train major vessel dimensions appear unchanged

from those used for the calculations however the ‘As Built’ dimensions of the E & W

Test Separators are greater and these vessels contribute a significant proportion of the

HP flare blowdown load. Any increase in HP flare blowdown load from this source can

be mitigated against the load incorporated for future equipment that remains

uninstalled.

No calculation was found to confirm compression train scrubber vessel dimensions

used for the blowdown/settleout calculations and therefore it was not possible to check

them against the ‘As Built’ vessels. This could have a significant effect on the LP flare

blowdown load. The missing calculation should be found and the calculation revised

to reflect the as built data.


32

Issue 34-005/3 - Were the real settle out pressures ever used?

As stated above, the validity of the blowdown/settleout calculations for the

compression train could not be confirmed. The ‘As Built’ calculation of these values

could have a significant effect on the LP flare blowdown load. The missing

calculations should be found and reviewed. If necessary the blowdown calculations

should be rerun and the results incorporated in the RABS update,

Issue 34-005/5 - Are vessel weights used reasonable?

Blowdown section weights stated in the calculation are based on vendor data for

vessels. Weight of pipework associated with major vessels appears to have been

estimated only. For systems that contain only pipework (e.g. manifold systems) major

pipework weight is calculated from the best available information at the time. It is

considered that a more accurate calculation of system weights would be unlikely to

have a significant effect on the blowdown loads (because of aggregate effects).

Issue 34-006/1 - HP Blowdown calculation higher than vendor aware of.

Radiation level for case is underestimated

Revision 06 of this calculation identified a HP flare blowdown load 5.8% greater than

the load used by the vendor for the flare radiation calculations. This increase is not a

concern at present because, as stated above, there is a significant allowance included

in the total HP flare blowdown load for future equipment. However, the increased HP

flare blowdown load should be incorporated into the updated RABS.

Issue 34-006/4 - Is constant rate blowdown a valid design method, i.e. not

according to API?

A constant blowdown rate (i.e. not reducing with time) was used for two items of low

pressure equipment, the LP separator and the LP fuel gas KO drum. Though not

normally valid, as they operate at low pressure and thus will not contribute a significant

proportion of the total LP flare blowdown load the calculation is considered acceptable

(see also Section 6.3.5).

Issue 34-006/5 - 'As Built' settleout pressure

As Issue 34-005/3 above.


33

Issue 34-010/1 – Was the calculation methodology sufficiently robust?

This calculation identified the minimum allowable temperatures that the process plant

could fall to during a process shutdown before potential problems could arise on

blowdown. The resulting cool-down temperatures were:

Hydrate Formation

Cooldown Temperature, oC 61.5

Resulting Blowdown Temperature, oC -30

Min Design Temp (-45 oC) Occurs in Flare


Resulting Blowdown Temperature, oC -45

The simulations used to generate these numbers were checked and they revealed that

the resultant blowdown temperatures were calculated using adiabatic flashes and not

using a blowdown model. The adiabatic flash assumes isenthalpic expansion, i.e. the

isentropic coefficient is 0, and results in higher downstream temperatures.

A blowdown model was run for the ‘Min Design Temp (-45 oC) Occurs in Flare’ case,

depressurising from the same settle out pressure and a temperature of 50 oC and gave

the following results:

Min Design Temp (-45 oC) Occurs in Flare – Blowdown Model


Resulting Blowdown Temperature, oC -59 (Isentropic Coefficient = 0.5)

Resulting Blowdown Temperature, oC -63 (Isentropic Coefficient = 0.7)

The results show that the temperature of the equipment must not be allowed to fall to

the level as originally calculated (50 oC) and more realistically blowdown should be

initiated at around 70 oC.


34

This calculation does not give any conclusions on the allowable delay before

blowdown should be initiated, this is addressed in calculation 34-060 / B for the

Injection Compressor only. 060 / B concludes that for the original model basis there is

a huge spread of allowable delay periods depending upon environmental conditions

and whether insulation is installed. The results of calculation 060 / B are given below:

Cooldown temp at which hydrate formation occurs in LP flare system on section blowdown = 61.5 C

Hold time for cooldown temperature to reach 61.5 C, hr: 3.4 (No insulation, natural convection)

0.96 (No insulation, forced convection)

20 (1" insulation, natural convection)

28 (1.5" insulation, forced convection)

Cooldown temp at which minimum design temperature occurs in LP flare system on section blowdown = 50 C

Hold time for cooldown temperature to reach 50 C, hr: 4.5 (No insulation, natural convection)




Current platform design philosophy is to depressurise after 1-2 hours. Given this large

spread the difference on calculation of the minimum allowable cooldown may not have

a significant effect on the delay allowed before blowdown is initiated. As no firm

conclusions were made in this calculation and given the problems in the input data this

whole issue should be revisited and re-evaluated using ‘As Built’ / operating equipment

and environmental data due to the possible adverse effects on the platform should the

minimum temperatures defined above be achieved. Once the new calculations had

been completed alarms could be added to the affected equipment (e.g. the gas

injection manifold) which would warn that blowdown was necessary. Allowing the

temperature to fall below this point would lead to excessively low temperatures and the

potential for flare pipework failure through embrittlement.

Issue 34-010/2 - Should 'troubleshooting' methanol injection points be

incorporated?

They actually are installed. Therefore no further concern.


35

Issue 34-011/1 – The individual well design rate has changed. What are the

implications for the platform?

The individual well rates have changed since design rendering the related calculations

obsolete. The new well rates need to be included in the RABS revision. Two aspects

will need to be addressed:

A final decision regarding the number of wells which fail to shut in, based on the

lower expected number of more prolific wells, which need to be designed for.

The maximum design well rate.

To follow from the above will require a modified procedure to be developed which

caters for the reduced time period available before the HP flare KO drum overfills

(which will happen in less than 10 minutes should relief occur at the higher well rates).

Issue 34-015/1 – Relief & Blowdown Study Report Rev C1 non-concurrent

maximum allowable LP and HP Flare loads are 110,874 kg/h

and 244,897 kg/h respectively. Rates used in these

calculations exceed design.

The maximum allowable independent LP and HP flare loads used in these calculations

to determine maximum allowable control valve CV for the blowby cases are greater

than the quoted figures in the Relief & Blowdown Study Report Rev C1 (i.e. 119,324

kg/h (LP) and 274,878 kg/h (HP) respectively). However the actual installed control

valve CVs are less than the calculated maximum. Therefore the system design rates

(244,897 and 110,874 kg/h) should not be exceeded. For consistency, the MP and LP

Separators relief valves should be checked against the installed control valve CVs.

Issue 34-015/2 - Is considering only one control valve fails open for the gas

blowby case when 2 are installed in parallel realistic / allowable

even with the provision of independent transmitters and

controllers?

This issue relates to the ability of only one of the LCVs to fail open. The level control

systems in question have independent transmitters and the equivalence in DCS terms

of duplicated controllers. This has been confirmed by HMDC studies. No further

action is required.


36

Issue 34-061/1 - Annulus rupture case had the potential to be the defining case

for the HP flare system (depending on installed choke valve CV).


The RABS states that gas lift to Hibernia wells is no longer required. If it becomes

necessary in the future then it will be provided by a method which will eliminate the

need to design for annulus rupture.

The situation with the Avalon wells is less clear. The RABS states that confirmation

was required during detailed design of the subsea facilities that annulus rupture need

not be considered as a relief case.

The need, or lack of, for design of the relief systems for annulus relief in both Hibernia

and Avalon wells should be confirmed.

Issue 34-022/1 – Calculated maximum pressure at PSV discharge exceeds value

on PSV datasheet Rev C1

Results of this calculation were taken into account on ‘As-Built’ PSV datasheet,

therefore no concern.

Issue 34-022/2 – Effect of increased production / production fluid GOR

The maximum associated gas capacity of the platform is governed by the capacity of

the compressors. This effectively set the required size of the HP separator relief

valves. Therefore even though the production GOR changes the relief valve should

have sufficient capacity whilst the compressors are able to take the gas. This link

should be made clear in the updated RABS.

Elsewhere in the system GOR has very little effect on the defining relief cases as

these are set by physical characteristics of installed valves, i.e. the gas blowby cases.

Should the physical characteristics of either of the above change, i.e. through

rewheeling a compressor stage, or through the use of larger control valve trims, the

calculations should be revisited.

Issue 34-023/1 – Calculated maximum pressure at spill-off valve discharge

exceeds value on control valve datasheet Rev C1

The difference in calculated and datasheet valve discharge pressure is small (146

kPa) and is insufficient to affect sizing. Furthermore, the installed CV is greater than

that required for the maximum flow (500 compared to 376 calculated for design

flowrate). Therefore the valve should be able to do its design duty.


37

Issue 34-023/2 – Is case where valve fails fully open is considered?

It appears this case was not considered (or, at least, not in the calculations we have

seen). This comment appears to be true for all the spillover valve cases. The

consequences of failed open spillover valves should form a section in the updated

RABS.

Returning to this particular case, the normal operation of this valve directs

245,000 kg/h to the HP flare (when the compressors shut down). This is the same

flowrate as the design maximum for the HP flare. The calculated valve CV for this flow

is 376 and the control valve installed CV is 500 therefore if the valve were to be sent

wide open, for any reason, there would be an instantaneous flowrate of around

326,000 kg/h directed to the HP flare, which is well above the HP flare design figure.

We are also aware that this valve trim has recently been replaced with a 550 CV trim

making the potential overshoot worse. The consequences of this relief case, such as

thermal radiation impingement on the platform, together with remedial measures to

limit the peak should be investigated further.

Issue 34-024/1 - Calculated maximum pressure at (MP separator) PSV discharge

exceeds value on PSV datasheet Rev C2

Results of this calculation taken into account on ‘As-Built’ PSV datasheet, therefore no

concern.

Issue 34-025/1 - Calculated maximum pressure at (3rd stage compressor) PSV

discharge exceeds value on PSV datasheet Rev C2

‘As-Built’ PSV datasheet retains original back pressure of 500 kPa(g) max. However

the PSV set pressure is 25,500 kPa(g) and the minor increase in back pressure will

have no effect on the PSV capacity.

Issue 34-025/2 - Is relief from both compressor trains a valid case?

We are aware of events which have caused relief valves to lift on both compressors

simultaneously. A modification project was included to avoid this occurrence. The

project should be reviewed for its capability to prevent this case and a description

should be included in the updated RABS.

Issue 34-026/1 - Calculated maximum pressure at PSV discharge exceeds value

on PSV datasheet Rev C2

‘As-Built’ PSV datasheet retains original back pressure of 500 kPa(g) max. However

as the minimum PSV set pressure is 45,500 kPa(g) the minor increase in back

pressure will have no effect on the PSV capacity.


38

Issue 34-027/1 – Was failed open MP spillover control valve considered?

The normal operation of this valve directs 94,500 kg/h to the HP flare. The calculated

valve CV for this flow is 522 and the control valve installed CV is 600 therefore if the

valve were to be sent wide open, for any reason, there would be an instantaneous

flowrate of somewhat less than 109,000 kg/h directed to the HP flare. This flowrate is

considerably lower than the HP flare design maximum so is tolerable. There may be

significant noise associated with this case.

Issue 34-028/1 – Is case where west test separator spillover valve fails fully open

considered?

The normal operation of this valve directs 58,800 kg/h to the HP flare. By inspection, if

the valve were to be sent wide open, for any reason, the instantaneous flow to the flare

would not exceed the HP flare design maximum so is tolerable. There may be


Issue 34-030/1 – Is case where east test separator valve fails fully open

considered?



would not exceed the HP flare design maximum so is tolerable. There may be


Issue 34-/1 – Calculated maximum pressure at 1st stage compressor spill-off

valve discharge exceeds value on control valve datasheet Rev C1

The difference in calculated and datasheet valve discharge pressure is small (50 kPa)

compared to the upstream pressure and the installed CV is greater than that required

for the maximum flow (320 compared to 231 calculated for design flowrate). The valve

will therefore easily pass the desired rate.

Issue 34-/2 – Is case where 1st stage compressor spillover valve fails fully open

considered?



would not exceed the HP flare design maximum rate. There may be significant noise

associated with this case.


39

Issue 34-039/1 – Is case where LP separator spillover valve fails fully open

considered?

The normal operation of this valve directs 60,200 kg/h to the LP flare. The calculated

valve CV for this flow is 3097 and the control valve installed CV is 4145 therefore if the

valve were to be sent wide open, for any reason, there would be an instantaneous

flowrate of somewhat less than 80,600 kg/h directed to the LP flare. This flow is

considerably less than the LP flare design maximum flowrate. The case would be of

short duration as the LP separator pressure would immediately begin to fall, lowering

the rate experienced.

Issue 34-045/1 - There is no network analysis run with common HP Blowdown at

initial conditions

The calculations for total HP blowdown were all done on an individual basis only to

check that the velocity criterion was not exceeded in the laterals (a check for

excessive pressure drop was also made). No network analysis for total blowdown was

done. This was a valid approach at the time as it could be assumed that all blowdown

valves would be operating at sonic velocities and back-pressures that could restrict

flow would never be reached in the system. Given that HMDC are considering

modifications to the platform, the construction of a network model of the HP flare

system would be a useful exercise to determine the effects of any modifications

proposed.

Issue 34-045/2 - Consistency error in the blowdown flowrate from the gas

injection flowlines.

This calculation identifies 8 blowdown valves each with an initial blowdown rate of

2250 kg/h (the blowdown simulation gives actual rate is 2233 kg/h). The total

blowdown rate for all GI lines given in calc 34-006/A is 8932 kg/h indicating that the

HP flare system is designed to accommodate four GI wells. If more than four GI wells

are installed in the future, calculation 006 should be revisited.

Issue 34-042/1 – Total blowdown rate (initial rate) used in calc less than that in

Relief & Blowdown Study Report (89,601 kg/h)

This calculation uses 86,709 kg/h for total LP blowdown initial rate, less than that

stated in Relief & Blowdown Study Report (89,601 kg/h). The lower blowdown load is

in fact a more up to date figure as stated in calculation 34-006 / A Rev 06. The

decreased LP flare blowdown load should be incorporated into an updated Relief &



40

Issue 34-033/1 – Assumption that the header is at zero pressure (i.e. that this is a

singular event not coincident with any other releases)

To calculate the PSV back pressure for these valves, it is assumed that the header the

lateral relieves to is at zero pressure. This will not be true for systems which do not

blowdown. During a local fire the affecting these systems the blowdown system will be

activated. A back pressure only slightly higher than that calculated will be greater than

10% of the PSV set pressure of 700 kPa(g). API RP520 only allows a maximum back

pressure of 10% of set pressure for this type of valve (conventional). The valve

selection therefore needs to be reviewed and resized as necessary. This will require a

flare network model to be constructed.

In a number of areas this same inconsistency exists and the valve selection should be

reviewed similarly.

Issue 34-036/1 – Inconsistency on (injection compressor PSV) datasheet

between accumulation and 'Max Relieving Pressure' (should be

10%)

Rev C2 PSV-7326 datasheet 'Max Relieving Pressure' is 121% of set pressure

therefore inconsistent with stated 10% accumulation. The ‘As Built’ datasheet shows

accumulation at 10% therefore no concern.

Issue 34-037/1 – Calculated back pressure (for 0152A/B) greater than specified

on datasheet - calc considers this OK as less than 10% of set

pressure

The maximum calculated PSV back pressure is 84 kPa(g); higher than the ‘As Built’

datasheet, which shows 1-35 kPa(g), but less than 10% of the PSV set pressure of

1380 kPa(g). API RP520 allows a maximum back pressure of 10% of set pressure for

conventional valves therefore there is no concern.

Issue 34-046/1 – 'As Built' P&IDs show bursting discs installed in this service

(calc considers PSVs) therefore calc is no longer valid

There is no replacement calculation for the installed bursting discs. The bursting disk

calculations should be reviewed to identify implications for the flare system.

Issue 34-050/1 – Rev C2 PSV datasheet states set pressure = 8200 kPa(g), 'As

Built' P&ID shows set pressure = 7000 kPa(g)

‘As Built’ datasheet has 33-PSV-7200 set pressure of 8200 kPa(g). As the item of

equipment the PSV is protecting (D-3303A) has a design pressure of 8200 kPa(g) the

error is on the P&ID. The P&ID should be corrected at the next revision. This also

applies to 33-PSV-7226 protecting D-3303B.


41

Issue 34-054/1 – Rev C2 (HP manifold) PSV datasheet states set pressure =

34,400 kPa(g), 'As Built' P&ID shows set pressure = 34,100

kPa(g)

The ‘As Built’ datasheet has 31-PSV-7042A/B set pressure of 34,100 kPa(g) therefore

the P&ID is probably correct and the design pressure revised downwards since the

Rev C2 PSV datasheet issued. No changes are therefore required.

Issue 34-057/1 – Calculated back pressure exceeds that specified on datasheet

for both (recirculation heater) PSVs

The maximum calculated back pressure for each PSV exceeds that stated on the ‘As

Built’ datasheet, which shows 1-35 kPa(g). However the calculated back pressure is

still less than 10% of each PSV set pressure. API RP520 allows a maximum back

pressure of 10% of set pressure for conventional valves therefore there is no concern.

Issue 34-059/1 – Accuracy of calculations using ESI instead of INPLANT

ESI has been used extensively in the flare calculations. This comparison calculation

between ESI and SIMSCI’s hydraulic simulator, INPLANT, showed that ESI gave

pressure drops 20% less than INPLANT. There appears that nothing was done to

recheck the calculations at the time. Assuming that INPLANT is more accurate (as it is

more rigorous) then a 20% difference on pressure drop calculation is significant. All

calculations using ESI should be revisited.


42

5.3 Results of the Technical Audit – Relief Valve Sizing Calculations

In Appendix I the full results of the audit are given. The detailed tables that follow

identify a number of issues to be dealt with which concern the accuracy or soundness

of the design conclusions.

Table 5.5 Technical Audit Issues – Relief Valve Sizing Calculations

Calculation Issue

Number Title Rev Date Number Description

31.35 Relief Valve Calculations - HP Separator

C1 Nov-91 31.35/1 Does 2 phase relief case become the governing case if the calculation new calculation method given in API RP520, Seventh Edition used?

31.35/2 Flare network analysis for 2 phase case

(Calc 34-064 / G) used total load = 252,372

kg/h (40,000 bpd).

31.35/3 Relief & Blowdown Study Report Rev C1

states HP Separator Blocked Outlet

(Vapour) relief load is 244,897 kg/h.

31.35/4 The two phase calculation feed vapour /

liquid split was abnormally low.

31.35/5 Methodological problem in calculation

(compared to API RP520 Sixth Edition).

The wrong effective pressure was for the

V/L split and property conditions.

31.36 Relief Valve Calculations - MP Separator

C1 Nov-91 31.36/1 Does 2 phase relief case become the

governing case if the calculation new

calculation method given in API RP520,

Seventh Edition used?

31.36/2 Are 2 x 50% LCVs sufficiently independent?

31.36/3 Methodological problem in calculation

(compared to API RP520 Sixth Edition).

The wrong pressure was used to generate

the vapour amount and properties.

31.36/4 The two phase calculation feed vapour /

liquid split was abnormally low.

31.36/5 Calculation subsequently superseded but

no indication that calculation was

subsequently corrected.

31.36/6 The gas blowby cases are methodologically

flawed.

31.36/7 There is an error in the gas rate calculated

by the test separator gas blowby case.


43

Table 5.3 Technical Audit Issues – Relief Valve Calculations (Cont.)

Calculation Issue

Number Title Rev Date Number Description

31.37 Relief Valve Calculations - LP

Separator

C0 27-Nov-91 31.37/1 Is it possible for the Test Separator


Separator when operating in high pressure

mode?


See also 31.36/6

31.38 Inlet Line Size Checking for Relief

Valves

05-Dec-91 31.38/1 Inlet line sizes should have been

recalculated using 'Final' relief data and

isometrics.

31.42 HP/MP/LP Separators PSV Inlet Line

Sizing

02-Jun-92 31.42/1 Pressure drop to HP Separator relief valves

has not been calculated using maximum

relieving capacity of valves

See also 31.43/1 & 31.43/2

31.43 Gas Blowby (Checking Capacity of

Downstream System for Gas Blowby

from HP to MP Separator and MP to

LP Separator)

22-Nov-92 31.43/1 This calculation considers both upstream

LCVs fail open simultaneously. This

scenario is not considered in the Relief &

Blowdown Study Report Rev C1 (or in any

other calculations reviewed), nor is the

platform designed for its affects.

31.43/2 This calculation considers both upstream

LCVs fail open simultaneously. This

scenario is not considered in the Relief &

Blowdown Study Report Rev C1 (or in any

other calculations reviewed), nor is the

platform designed for its affects.

31.43/3 The calculation identifies the failure of the

spillover valve (open) could lead to a relief

rate which is higher than the current design.

5.3.1 Technical Audit Issue Discussion – Relief Valve Calculations

The following describes the technical audit issues identified in Table 5.3. In this case

the issues are grouped by relief valve and then under broad issue headings.


44

5.3.1.1 HP Separator Relief Valves

Two phase case

A series of issues raised by the technical audit can be grouped together under this

issue, i.e.

Issue 31.35/1 - Does two phase relief case become the governing case if the

calculation new calculation method given in API RP520, Seventh

Edition used?

Issue 31.35/2 - Flare network analysis for 2 phase case (Calc 34-064 / G) used

total load = 252,372 kg/h (40,000 bpd).

Issue 31.35/4 - The two phase calculation feed vapour / liquid split was

abnormally low.

Issue 31.35/5 - Methodological problem in calculation (compared to API RP520

Sixth Edition). The wrong pressure was used to generate the

vapour amount and properties.

The calculation of the relief valve area for the two-phase case underestimates the area

required. There was a combination of inconsistency, probably incorrect simulation

compositions and a flaw in methodology (compared to API RP 520 Sixth Edition) which

together would underestimate the orifice area required. However, because of the new

sizing method we are obliged now to use (see Section 6.5) and because of well rate

considerations these problems will naturally be corrected in the new calculation that

will be required.

So returning to 31.35/1, the most important of these considerations; the relief valve

orifice area required for the original two phase relief case based on the new Leung

omega API RP520 calculation method is 8.21 in2. This compares with the originally

(incorrectly) calculated value of 3.19 in2 for the same relief scenario. Obviously the

new method of calculation has a significant effect on required orifice area for relieving

two phase flow. In this particular case, the governing case for the relief valve was

blocked outlet – vapour relief only, which required a minimum orifice area of 9.38 in2

whereas the actual installed orifice area is 11.05 in2. On the face of it, therefore, no

hardware modifications are required.

However, we are aware that the maximum single well rate is now well in excess of the

originally considered 40 kbopd case (which represented one large and one medium

well failing to shut in). Therefore the relief case must be rigorously recalculated.


45

Using the new sizing method the maximum safe well rate for a single well failing to

shut in is approximately 54 kbopd. Measures should be taken to limit the maximum

well rate to this value, and to less than this value if 2 wells failing to shut in becomes

the selected basis.

Vapour only case

Issue 31.35/3 - Relief & Blowdown Study Report Rev C1 states HP Separator

Blocked Outlet (Vapour) relief load is 244,897 kg/h.

The rate used in the calculation and to purchase the relief valve was 227,649 kg/h.

This appears to be confused in the RABS with the normal maximum associated gas

rate when the spillover valve is open during a compressor shut down. This

inconsistency should be corrected in the updated RABS.

5.3.1.2 MP Separator Relief Valves

Two phase case

Issue 31.36/1 - Does two phase relief case become the governing case if the new

calculation method given in API RP520, Seventh Edition is used?

Issue 31.36/3 - Methodological problem in calculation (compared to API RP520

Sixth Edition). The wrong pressure was used to generate the

vapour amount and properties.

Issue 31.36/4 - The two phase calculation feed vapour / liquid split was

abnormally low.

The same description as above (Section 5.3.1.1) is equally valid here (although the

orifice areas are different). This case must be calculated rigorously recalculated.

Issue 31.36/2 - Are 2 x 50% LCVs sufficiently independent?

This issue relates to the ability of only one of the LCVs to fail open. The level control

systems in question have independent transmitters and the equivalence in DCS terms

of duplicated controllers. This has been confirmed by HMDC studies. No further

action is required. See also below.

Issue 31.36/5 - Calculation subsequently superseded but no indication that

calculation was subsequently corrected.

The gas blowby rate was reduced in January 93 (because the valve CV was reduced

from 400 to 350). However the valve selection remained unchanged. The required

orifice area is therefore higher than it needed to be.


46

Issue 31.36/6 - Are the gas blowby cases methodologically flawed?

It is arguable that the gas blowby cases are methodologically flawed. The reasons for

this are given below:

API RP520 does not allow credit for control valves in their normal position to move

to reduce the relief rate, although the requirements are rather vague. Thus there

may be a case for considering, irrespective of the control system, the worst case

as one valve failed open and one in its normal position. However, as there is also

a ESD/PSD between the systems we believe this would a very harsh case to

consider.

The 100% gas case is normally considered only in a shutdown situation (otherwise

a high component of the feed to the separator must pass with the gas to low

pressure system and be relieved). Therefore, in a shutdown situation (where the

liquid has dumped for some reason) a calculation method similar to settleout is

normally used, i.e. the volume of gas lost from the higher pressure side to raise the

pressure of the lower pressure side to the relief pressure is removed from the

effective relief driving force. Also in this case the failure open of one valve is more

likely to be acceptable because the likelihood of the both LCVs and the isolating

ESVs failing is very low

Should a LCV fail open during normal production then the blowby fluid is both

vapour and liquid (which will reduce the effective blowby volume rate) and also the

normal positions of the downstream control valves could be taken into account

thereby dramatically reducing the apparent relief rate.

The above assumes the QRA did not identify the possibility of both valves failing open

which we understand to be the case.

The net effect of the above is to suggest the relief rates designed for are higher than

they need be. A note could be added to the updated RABS to reflect this.

Issue 31.36/7 - There is an error in the gas rate calculated in the test separator

gas blowby case.

The calculation of the gas blowby rate from the test separators uses a subcritical

formula for the level control valve even though the calculation shows that the flow is

critical. By inspection, the error will not affect the sizing of the relief valve as relief rate

for this case is considerably less than for the governing case.


47

5.3.1.3 LP Separator

Issue 31.37/1 - Is it possible for the Test Separator manifold to be connected to

the LP Separator when operating in high pressure mode?

The test separators are able to operate in two modes, high pressure and low pressure.

When operating in HP mode they are connected to the MP Separator and when in LP

mode to the LP Separator. This calculation considers that the test separators are in

LP mode but the reliability of the installed precautions / interlocks preventing

connection of the test separators in HP mode to the LP Separator is not apparent. As

the test separators can operate at 4240 kpa the potential gas blowby rate to the LP

Separator, if incorrectly lined up, would be significant. This relief scenario should be

investigated further.

Issue 31.37/2 - Are 2 x 50% LCVs sufficiently independent?

See Issue 31.36/2 above.

5.3.1.4 Miscellaneous

Issue 31.38/1 - Inlet line sizes should have been recalculated using 'Final' relief

data and isometrics.

This calculation was done with preliminary data and has obviously been revised as

many PSV inlet line sizes shown on the ‘As Built’ P&IDs are different to those

calculated here. The inlet line sizes should be checked against ‘As Built’ data and

isometrics.

Issue 31.42/1 - Pressure drop to HP Separator relief valves has not been

calculated using maximum relieving capacity of valves.

The relief valve inlet line size has been calculated using the calculated governing case

relief rate of 227,649 kg/h. API RP520 Part II states that the inlet line size should be

calculated using the ‘maximum rated capacity’ of the installed relief valve which in this

case is 262,161 kg/h. It is not expected that the inconsistency will have a significant

effect on the inlet line size as a margin of 20% was applied at the time. (See also

Issue 31.38/1 above).

Issue 31.43/1 and 2 - This calculation considers both upstream LCVs fail open

simultaneously. This scenario is not considered in the Relief &

Blowdown Study Report Rev C1 (or in any other calculations

reviewed), nor is the platform designed for its affects.

In view of the description in 31.36/6 this case does not appear feasible. The notes

attached to the calculations should have said so.


48

Issue 31.43/3 - The calculation identifies the failure of the spillover valve (open)

could lead to a relief rate which is higher than the current design.

The calculation identifies that the LP Separator spillover valve, if it failed fully open,

could generate a flowrate of 121,035 kg/h in the LP flare. This is greater than the

current design LP flare capacity of 110,874 kg/h. The maximum flowrate which could

be sent to the LP flare system under this scenario should be investigated for the

current operation.

5.4 Technical Audit Conclusion Summary

In tabular form, the following summarises the actions required to be undertaken in

Stage 2.

Table 5.6 Technical Audit Conclusion Summary (System 34)

Number

34-

Title Number

34-

Description Action




calculations

005/2 Jet fire scenario not taken into account

for the design of the blowdown system

Incorporate jet fire calculations

and update RABS accordingly


used?




005/4 Were fire areas used for total blowdown

rate?

No further action

005/5 Are vessel weights used reasonable? No further action




Update RABS.





See also 34.010/1.

006/3 Is design case too extreme? Select start pressure basis and

update RABS.

006/4 Is constant rate blowdown a valid design

method, i.e. not according to API?

No further action unless staggered

blowdown becomes an issue

006/5 'As Built' settleout pressure See 005/3 above


cooldown before

hydrate formation &

minimum

temperatures



sections











010/2 Should 'troubleshooting' methanol

injection points be incorporated?

No further action


49

Number

34-

Title Number

34-

Description Action


KO Drum size







platform?






012 / A Review of LP flare KO

Drum size

See 34-011/1 See 34-011/1



Separator and MP to

LP Separator Blowby

Cases






design.




015/2 Is considering only one control valve

fails open for gas blowby case when 2

installed in parallel realistic / allowable

even with provision of independent

transmitters and controllers?

No further action required

060 / B Indicative Injection

Compressor

Cooldown Calculation

See 34-010/1 and 34-010/2 See 34-010/1 and 34-010/2

061 / B Simplistic Steady

State Preliminary

Review of the Annulus

Rupture Relief

Flowrate

061/1 This case had the potential to be the

defining case for the HP flare system

(depending on installed choke valve CV)





- Max Relief Case)

022/1 Calculated maximum pressure at PSV

discharge exceeds value on PSV

datasheet Rev C1






capacity.

023 / C HP Separator Max

Spill-off Case -

Network Analysis

023/1 Calculated maximum pressure at spill-off

valve discharge exceeds value on

control valve datasheet Rev C1

No further action required.


considered?

Recalculate case. Assess

measures for reducing the peak

load during failure. Implement

modification project.

See also 34-022/2 See also 34-022/2

024 / C MP Separator Max


Analysis



datasheet Rev C2



50

Number

34-

Title Number

34-

Description Action

025 / C 3rd Stage

Compressor Max


Analysis



datasheet Rev C2 - Check for later

revisions










026 / C Injection Compressor

Max Relief Case -

Network Analysis



datasheet Rev C2


See also 34-025/1& 34-025/2 See also 34-025/1& 34-025/2

027 / C MP Separator Max

Spill-off Case -

Network Analysis

027/1 Was failed open control valve

considered?



load during failure if necessary.

See also 34-022/2

028 / D West Test Separator

Max Spill-off Case -

Network Analysis


considered.



029 / D West Test Separator

Max Relief Case -

Network Analysis

See 34-022/2 See 34-022/2

030 / D East Test Separator

Max Spill-off Case -

Network Analysis

030/1 Is case where valve fails fully open is

considered?



031 / D East Test Separator

Max Relief Case -

Network Analysis

See 34-022/2 See 34-022/2

34- / E 1st Stage Compressor

Spill-off Case -

Network Analysis

34-/1 Calculated maximum pressure at spill-off

valve discharge exceeds value on

control valve datasheet Rev C1


34-/2 Is case where valve fails fully open

considered?


045 / E Total HP Blowdown

Initial Conditions

(Checks blowdown

line sizes for

individual system

blowdowns)

045/1 There is no network analysis run with

common HP Blowdown at initial

conditions

Consider constructing a HP flare

network model to assess future

modification projects against.

045/2 Consistency error in the number and

flows in the gas injection flowlines

Add a note to the RABS clarifying

the injection manifold rate basis.

039 / F LP Separator Max

Spill-off Case -

Network Analysis


considered?



load if necessary.



51

Number

34-

Title Number

34-

Description Action

042 / F Total LP Blowdown -

Initial Conditions -

Network Analysis

042/1 Total blowdown rate (initial rate) used in

calc less than that in Relief & Blowdown

Study Report ( 89,601 kg/h)

Update RABS

042/2 Validity of staggering blowdown. Were

the systems sufficiently independent?

Perform safety analysis to

satisfactory standard to show the

vessels will not fail during jet fire

(including the A injection

compressor and components).

043 / F Injection Compressor

'A' Blowdown - Initial

Conditions - Network

Analysis

See 34-042/2 See 34-042/2


Separator Heaters

Simultaneous Fire

Relief - Network

Analysis




Construct a LP flare network




036 / G Injection Stage

Suction Scrubber PSV

- Network Analysis

036/1 Inconsistency on datasheet between

accumulation and 'Max Relieving

Pressure' (should be 10%)


037 / G HM & CM Expansion

Drums Simultaneous

Fire Relief Case -

Network Analysis

037/1 Calculated back pressure (for 0152A/B)

greater than specified on datasheet -

calc considers this OK as less than 10%

of set pressure

See also 34-033/1


044 / G Total LP Blowdown -

After 3 mins (stagger

point) - Network

Analysis

See 34-042/2 See 34-042/2

046 / G Fuel Gas Cooler /

Heater tube rupture


046/1 ''As Built' P&IDs show bursting discs in

this service (calc considers PSVs)

therefore calc is no longer valid

There is no replacement

calculation for the installed

bursting discs. The bursting disk

calculations should be reviewed to

identify implications for the flare

system.

050 / G 3rd Stage Suction

Scrubber A (D-3303A)

PSV Discharge Line

Size Confirmation

050/1 Rev C2 PSV datasheet states set

pressure = 8200 kPa(g), 'As Built' P&ID

shows set pressure = 7000 kPa(g)

P&ID set pressure error?

052 / G E-3301 Shell Side

PSV Discharge Line

Size Confirmation

See 34-033/1 See 34-033/1

053 / G E-3303B Shell Side

PSV Discharge Line

Size Confirmation

See 34-046/1 See 34-046/1


52

Number

34-

Title Number

34-

Description Action

054 / G HP Manifold Relief -

Network Analysis


pressure = 34,400 kPa(g), 'As Built'

P&ID shows set pressure = 34,100

kPa(g)


055 / G Simultaneous Fire

Relief Case from Z-

3701 A/B, Z-3702 A/B

& Z6202 A/B (pig

launchers and fuel

gas package)

See 34-033/1 See 34-033/1

057 / G E-3701 Shell & Tube

Side Simultaneous

Fire Relief Case - Line

Size Confirmation

057/1 Calculated back pressure exceeds that

specified on datasheet for both PSVs

See also 34-033/1

See 34-033/1

058 / G E-6201A/B Tube Side

Fire Relief Case

See also 34-033/1 See 34-033/1

059 / G Comparative Program

check of INPLANT

Single Phase

Simulation vs ESI

059/1 Accuracy of calculations using ESI

instead of INPLANT

Revisit ESI calculations and

replace as necessary


53

Table 5.7 Technical Audit Conclusion Summary (System 31)


31.35 Relief Valve Calculations - HP Separator

31.35/1 Does 2 phase relief case become

the governing case if the calculation

new calculation method given in API

RP520, Seventh Edition used?

Replace two phase sizing case

with new method. Use finalised

maximum well design rates.

Implement modification project as

necessary.

31.35/2 Flare network analysis for 2 phase

case (Calc 34-064 / G) used total

load = 252,372 kg/h (40,000 bpd).

See 31.35/1

31.35/3 Relief & Blowdown Study Report

Rev C1 states HP Separator

Blocked Outlet (Vapour) relief load is

244,897 kg/h.

Correct inconsistency in RABS

update.

31.35/4 The two phase calculation feed

vapour / liquid split was abnormally

low.

See 31.35/1

31.35/5 Methodological problem in

calculation (compared to API RP520

Sixth Edition). The wrong effective

pressure was for the V/L split and

property conditions.

See 31.35/1


31.36/1 Does 2 phase relief case become

the governing case if the calculation

new calculation method given in API

RP520, Seventh Edition used?

See 31.35/1

31.36/2 Are 2 x 50% LCVs sufficiently

independent?


31.36/3 Methodological problem in

calculation (compared to API RP520

Sixth Edition). The wrong pressure

was used to generate the vapour

amount and properties.

See 31.35/1

31.36/4 The two phase calculation feed

vapour / liquid split was abnormally

low.

See 31.35/1

31.36/5 Calculation subsequently

superseded but no indication that

calculation was subsequently

corrected.


31.36/6 Are the gas blowby cases are

methodologically flawed?

Add note to RABS update

31.36/7 There is an error in the gas rate

calculated by the test separator gas

blowby case.


31.37 Relief Valve

Calculations - LP

Separator




pressure mode?




this.

31.37/2 Are 2 x 50% LCVs sufficiently

independent?

See 31.36/2

See also 31.36/6 See also 31.36/6


54


31.38 Inlet Line Size

Checking for Relief

Valves

31.38/1 Inlet line sizes should have been

recalculated using 'Final' relief data

and isometrics.

Check / redo inlet line sizing

calculations as necessary.

31.42 HP/MP/LP Separators

PSV Inlet Line Sizing

31.42/1 Pressure drop to HP Separator relief

valves has not been calculated using

maximum relieving capacity of

valves

See 31.38/1

See also 31.43/1 & 31.43/2 See also 31.43/1 & 31.43/2

31.43 Gas Blowby

(Checking Capacity of

Downstream System

for Gas Blowby from

HP to MP Separator

and MP to LP

Separator)

31.43/1&2 This calculation considers both

upstream LCVs fail open

simultaneously. This scenario is not

considered in the Relief & Blowdown

Study Report Rev C1 (or in any

other calculations reviewed), nor is

the platform designed for its affects.


31.43/3 The calculation identifies the failure

of the spillover valve (open) could

lead to a relief rate which is higher

than the current design.



load during failure. Implement

modification project. Same action

as 34.039/1.


55

6.0 CHALLENGE PROCESS

6.1 Introduction

Before commencing the challenge process the key legislative aspects are introduced

6.1.1 The Principles of the Legislation

The general requirements of HSW legislation are neatly summarised in the following

extract. Here we use an interpretation supplied by a representative of the relevant UK

government department. We believe the requirements of Canadian HSW legislation to

be very similar:

As a duty holder under HSW legislation, you have a continuing duty to ensure the H&S

of employees and other persons who may be affected by the way in which you

undertake your business. The legislative regime sets a goal for duty holders to do all

that is reasonably practicable. In some areas technological change is so fast that

standards of compliance which would have been acceptable 10 years ago, are no

longer satisfactory. However the advantage of goal setting is that it keeps pace with

technological change, but also allows you to develop solutions which are better

aligned with the risks in your workplace.

So you will always, and continuously, have to keep an eye on new codes, standards,

good industry practice, etc., to ensure you are doing enough to satisfy the law. Where

it is reasonably practicable to do so, changes should be made. Nevertheless HSE

would accept that for existing installations it may be less reasonably practicable to

make a change, than is the case for a new installation - it is a judgement call which the

law requires you to make (and be able to justify, if or when challenged).

This sets the general principles by which retrospective change can be considered on

Hibernia. This advice clearly states that new codes and practices should be applied to

the facility unless it can be shown to be unreasonable.

6.1.2 Relevant Canadian Legislation

The relevant legislative regulations for Hibernia are those issued under the Canada -

Newfoundland Atlantic Accord Implementation Act. The key regulations which have

relevance for the flare system are as follows:

Newfoundland Offshore Petroleum Installations Regulations

During the design phase the relevant version of the regulations was the 1991 draft.

These were revised once again in draft form in 1993. The regulations were finally

registered on the 21 February, 1995.

During this time there was no substantive change to the documents in relation to

the design of the flare system.


56

The key extracts from the 1995 regulations are given below:

Gas Release System

17. (1) In this section, “gas release system” means a system for releasing gas and

combustible liquid from an installation and includes a flare system, a pressure relief

system, a depressurizing system and a cold vent system.

(2) Every gas release system shall be designed and located, taking into account

the amounts of combustibles to be released, the prevailing winds, the location of

other equipment and facilities…, so that when the system is in operating it will not

damage the installation…, or injure any person.

(3) Every gas release system shall be designed and installed in accordance with

(a) American Petroleum Institute RP 520, Recommended Practice for the Design

and Installation of Pressure-Relieving Systems in Refineries;

(b) American Petroleum Institute RP 521, Guide for Pressure -Relieving and

Depressuring Systems;

(c) American Petroleum Institute Standard 526, Flanged Steel Safety-Relief

Valves;

(d) American Petroleum Institute Standard 527, Seat Tightness of Pressure Relief

Valves; and

(e) American Petroleum Institute Standard 2000, Venting Atmospheric and Low-

Pressure Storage Tanks.

(4) Every gas release system shall be designed and constructed to ensure that

oxygen cannot enter the system during normal operation.

(5)…

(7) With the exception of water, any liquid that cannot be safely and reliably burned

at the flare tip of a gas release system shall be removed from the gas before it

enters the flare…

(9) Every gas release system shall be designed and installed so that, taking into

account the prevailing wind conditions, the maximum radiation on areas where

personnel may be located , from the automatically ignited flame of a flare or vent,

will be

(a) 6.3 kW/m2, where the period of exposure will not be greater than one minute;

(b) 4.72 kW/m2, where the period of exposure will be greater than one minute but

not greater than one hour; and


57

(c) 1.9 kW/m2, where the period of exposure will be greater than one hour.

The remainder of the regulations describe requirements relating to ventilation, the

selection of electrical components, escape routes, emergency control systems etc.

Petroleum Occupational Safety and Health Regulations - Newfoundland, Draft

November 1989.

These regulations are referenced in the RABS. Their only influence on the flare

system design appears to be to ensure the sound levels are acceptable:

i.e.

85 dB but no more than 90 dB for 8 hours exposure

102 dB but no more than 104 dB for 1 hours exposure

etc.

The other regulations covering certificates of fitness, drilling, and diving have no

significant influence on the design of the flare system.

6.1.3 Applying the Legislation

With the legislation in mind, the remainder of this section is organised to deal first with

the larger issues, which may affect a number of design characteristics of the flare

system. Subsequently the lesser aspects are dealt with sequentially as required.


58

6.2 Jet Fire

6.2.1 Requirements of the relevant regulations, design codes and practices when

Hibernia was designed

6.2.1.1 Canadian Legislation

We are unaware of any explicit references in Canadian legislation regarding this

particular hazard.

6.2.1.2 Mobil Engineering Guide (EGS 661-1990)

We have been unable to source a copy of the 1990 version of the above so instead

are required to interpret between the 1985 version and the Draft 1991 version. In this

case no interpretation is required, as both versions are silent on the subject of jet

flame and the design requirements to mitigate against its effects. The guides,

however, do mention the maximum time allowed for depressuring is 2 minutes per

3 mm of vessel wall thickness, but shall not be less than 6 minutes. Where

depressuring is impractical the use of water sprays or insulation (both applied to Mobil

specification) can be considered. Some of these latter methods could be read to take

some account of extreme fires (for example jet fire).

6.2.1.3 API 521 (Third Edition, November 1990)

The third edition is silent on the issue of jet fire impingement on vessels.

6.2.2 How Jet Fire Was Actually Handled During Design

This is a summary of how jet fire was handled in the design phase. The entire subject

of its consideration was based in the probabilistic safety related design path. Out of

necessity it is an abridged version as the component parts are numerously described

in the various project documentation. Here the aim is to capture the essence of the

process and its effect on the way Hibernia was designed.

This is described below.

6.2.2.1 Legislative Requirements

The Canadian draft Production and Conservation Regulations required the submission

and maintenance of a Safety Plan for the Hibernia facilities. Part of the Safety Plan

would comprise the Concept Safety Evaluation. This document would form the basis

for all the risk and hazard related studies on the Hibernia Project.


59

6.2.2.2 Concept Safety Evaluation (CSE)

The Concept Safety Evaluation (Reference 2) was the first attempt to identify and

quantify the major hazards connected with the Hibernia facility. The study also began

to define the terms Design Accidental Event (DAE) and Residual Accidental Event

(RAE) which would be used for the rest of the project. The definitions would, however,

gradually change over time but for the moment it is understood a DAE would only

affect those in the immediate vicinity of the accident whereas a RAE would potentially

affect the platform population.

Of key interest to the Flare Revalidation Study were the aspects identified regarding jet

fire.

The study explicitly recognised the potential for vessel rupture caused by short

duration jet fires in Module M10.

Deluge was considered ineffective in preventing escalation due to jet fire.

The times given for failure from jet fire impingement on various thicknesses of steel

were:

Table 6.8 Failure Times for Structures Engulfed in Jet and Pool Fires

Structure Jet Fire* (min) Pool Fire+ (min)

60mm Thickness Steel 12 30

25mm Thickness Steel 5 13

12mm Thickness Steel 2.5 6

5mm Thickness Steel 1 2.5

H120 Firewall 60 120

13mm Thickness Steel Tube Coated with Chartek Type III** >60

* Based on a jet fire radiation of 300 kW/m2

+ Based on a pool fire radiation of 150 kW/m2

** Based on research and field trials by Shell Thornton

(CSE Table 6.1)

The CSE also recognised the Key Safety Functions defined as follows:

the platform’s primary structure

the escape routes from the central parts of the platform to the Temporary Safe

Refuge

the Temporary Safe Refuge (TSR), including the central control room

the availability of the evacuation systems.


60

Added to this were Hibernia’s three-tier framework of risk acceptability:

For any single incident that might affect the key safety systems (more accurately

functions from the above), the risk level for the three-tiers are:

Intolerable: greater than 10-4 per year.

ALARP region: 10-4 to 10-5 per year.

Lower bound of acceptability: less than 10-5 per year

The CSE went on then to assess the risk to the Key Safety Functions using

consequence analysis and event trees.

The CSE demonstrated to a reasonable extent that the effects of jet fire and explosion

did not jeopardise the structural integrity of the platform or the availability of the

evacuation systems. In this it is implicit in the CSE that jet fire was considered a RAE

and more of a risk to structural impairment than explosion.

The CSE also indicated that jet fire impingement may cause rapid failure of

unprotected structures even if deluge systems are operating. This might be because

the intense heating raises the surface temperature above 100°C, prior to application of

water, preventing the formation of a protective liquid film.

To confirm the above the CSE made various recommendations for future work which

included the requirement to conduct a Fire Risk Assessment to review the impact of

fire on structural integrity of the H120 walls and the flare boom and the potential for

escalation in Module M10.

6.2.2.3 Fire Risk Assessment (FRA)

By the time the FRA (Reference 3) was commenced the HMDC Damage / Impairment

Criteria had been formalised. These can be found in Section 3.2.1:

Also outlined in the FRA were the details of the blowdown system. The system

considered was, in principle, the same as the system outlined in the Relief and

Blowdown Study Report and subsequently built (see Table 6.17 for more detail).

The FRA looked at the duration and flame lengths of jet fires with and without

blowdown. This is summarised below:

Table 6.9 HP Separator Jet Flame Length With and Without Blowdown

Hole Size Without Blowdown (m) With Blowdown (at the end of the 15

minutes) (m)

5 mm 9 5

50 mm 53 29


61

Wherever the possibility of a jet fire impinging on a structural member was identified

the FRA recommended the use of PFP and firewalls to ensure the impairment criteria

were satisfied. These aspects were studied in the Structural Passive Fire Protection

Analysis (Reference 7).

The analysis undertaken in the FRA identified the potential for escalation should a jet

fire impinge on a vessel, e.g.

“Fire water deluge will act to keep the equipment cool, but a jet fire impinging directly

on a vessel may cause localised heating and loss of wall strength…”

(FRA page 56)

“Operation of the blowdown system should not be viewed as evidence of satisfactory

vessel response, and may not prevent failure if the vessel is subject to high heat loads

such as jet flame impingement...”

(FRA page 57)

“A HP separator incident could escalate to the LP separator and vice versa. The

vessels will be provided with local deluge protection which will provide adequate

protection for incident thermal radiation…”

(Emphasis added. FRA page 70)

The emphasis is added to contrast against protection from jet flame impingement.

This was more clearly outlined in the team review (part of the consequence analysis):

“The LP separator is likely to fail due to jet flame impingement…”

(FRA page B.15)

This led to the recommendation to install kerbs to prevent spread of liquid spills or pool

fires.

Generally the problem of jet fire impingement on vessels was implicitly mentioned on a

number of levels in Module M10. Other jet fire consequence analyses identified the

problems of jet flame impingement on firewalls, the crude oil coolers, the flare boom

and the hydraulic panel on level 5.

One of the key conclusions of the FRA was less clear:

“The FRA considered a number of potential fire scenarios. In each scenario, except

possibly a large blowout, the active systems (isolation, blowdown, F&G detection and

protection) will limit the consequences and should prevent escalation.


62

Even if the active systems fail to operate, then the passive fire protection should

ensure that the Damage / Impairment Criteria are met…”

(FRA page 119)

The recommendations were to be studied further.

6.2.2.4 FRA Update

The FRA Update (Reference 4) revisited the key aspects of the FRA in relation to jet

fire and began to soften the conclusions. Some of the statements are included below:

“The KO drums are provided with deluge. This may not provide complete protection

against jet flame impingement…, but the duration would be short and failure is

unlikely…”

(FRA Update page 20)

“A jet flame from the gas scrubbers could impinge the MP separator, but it is unlikely

that it would be of sufficient duration to cause failure, provided the blowdown system

operates…”


A leak from either the HP or LP Separators could cause either a jet flame or a pool

fire…Escalation to the HP and LP Separator is unlikely provided that the deluge

operates and the vessels are blown down…”


A related aspect considered in the FRA Update was the use of PFP and particularly

Lloyds who stated no credit should be taken for any active fire systems when

considering the ability of PFP systems.

The study ended with the main FRA conclusions being considered valid.

6.2.2.5 Design Phase Risk Assessment of Potential Accidental Events (DPRA)

The DPRA (Reference 5) focused on the different types of accidental events. The

concept of contained events being DAE and uncontained events being RAE was

formally introduced.

The report also contained discussion of how the design was optimised so as to

prevent DAEs escalating into RAEs and impairing the main safety functions. The

document formalised the selection and differentiation between DAE and RAE. This is

shown below:


63

Figure 6.2 Method for Determining the Acceptability of DAEs and RAEs


64

The report was not completely clear as to which events were RAE or DAE. However

the following outlines our understanding:

Some hydrocarbon release events (assumed to be explosion, blowout and smoke)

are RAEs.

Otherwise, most fire hazards are DAEs. This includes jet fire impingement on main

structural members as this was mitigated against using PFP. Only in a few cases is

there potential for the fire scenarios to cause a RAE. These were identified as:

Fire damage to deck plate at el 114.000 allows fire damage to the underdeck

and collapse of structures within 2 hours.

Spread of fire from one wellhead to the another exceeds the capability of the fire

protection systems.

Failure of the isolation and blowdown systems to contain the inventory of

produced hydrocarbons and failure of the fire fighting systems to prevent

escalation for both topsides and utility shaft hydrocarbon releases. These are

“worst case” scenarios where virtually all the emergency systems have failed to

operate as intended. (An Emergency Systems Report would be prepared to look

into the possibility of these failures).

Smoke and heat effects could cause impairment of the Temporary Safe Refuge

(TSR) if the worst case circumstances occurred, e.g. the wind blows towards the

TSR, HVAC systems fails to detect smoke / gas or shutdown and doors and

penetrations are open.

Blowout combined with worst case weather conditions, causes impairment of the

TSR and evacuation systems.

Large fire on a hydrocarbon deck in the Utility shaft causing significant spalling of

the concrete walls and failure of the reinforcing bars.

All the above were subjected to further study and CBA (whose requirement that the

mitigation measure be undertaken if the cost was less than 10 times the yearly loss) to

show the risks were ALARP.

In the case of the failure of the safety systems the RAE was considered a DAE after

further study:

“In exceptional circumstances, individual components might be impaired. However,

the system’s redundancy was found to be adequate to perform the intended function”

(Reference 8)

“The Explosion Overpressure Risk Assessment… further assessed the effects of

explosions on the vent and blowdown systems and the structural design basis (one of

the main safety functions). Both were found to be acceptable.


65

“Thus it is concluded that the basis for the selection of DAE and RAEs, and the

assumptions relating to the adequacy of emergency systems preventing DAEs into

RAEs, are valid.

DPRA page 42

The remainder of the report explains the other types of event which are RAE. These

are summarised in the table overleaf:


66

Table 6.10 Summary of Risks of Impairment (as they relate to this study)

RESIDUAL ACCIDENTAL

EVENT

RISK OF IMPAIRMENT OF MAIN SAFETY FUNCTION

(per year)

CSE Estimate Detailed Design Risk

Hydrocarbon Fire Hazards:

Damage to Underdeck

Spread to Wellheads

Heat Impairment of TSR

Crude Oil Fire in Utility Shaft

-

-

-

-

2 x 10-6

1.5 x 10-6

1 x 10-6

3 x 10-7

Sub-total for Hydrocarbon Fire Hazards - 4.8 x 10-6

Explosion Hazards

M10

M20

-

-

-

1 x 10-6

1 x 10-6

Sub-total for Explosion Hazards - 2 x 10-6

Blowout 1 x 10-4 1.5 x 10-5

Smoke 1.7 x 10-4 (in QRA of TSR

Integrity report)2.1 x 10-5

Dropped Objects - -

Flooding of Utility Shaft - <1 x 10-6

External Events:

Iceberg Collision

Ship Collision

Helicopter Crash

Earthquake

Wave Slam

5 x 10-7

5 x 10-7

1 x 10-6

(Note 2)

-

5 x 10-7

5 x 10-7

Negligible (Note 1)

(Note 2)

5 x 10-5 (Note 3)

Sub-total for External Events 2 x 10-6 1 x 10-6

TOTAL 2.7 x 10-4 1.5 x 10-5

Notes

Note 1: Negligible risk of impairing main safety functions. Risk to occupants of helicopter could be 2.6 x 10-4 per year.

Individual risk will be lower (approximately half) because the same individual is not on every helicopter flight.

Note 2: Earthquake design return period is 2000 years (5 x 10-4 per year) but this does not lead to structural failure

nor pollution. Risk associated with hydrocarbon events caused by earthquake will be significantly less than

other causes.

Note 3: Risk of damaging one essential generator. Risk to main safety functions will be negligible.

Jet fires are mentioned no further in the document.


67

6.2.2.6 Design Phase Safety and Environmental Evaluation (DPSEE)

The preceding documents were summarised in the DPSEE (Reference 6). There

were no fundamental changes of relevance to the analysis herein.

It is relevant to mention the process areas in M10 where jet fire impingement was

explicitly considered and the mitigating reasons given:

Flare system - It was identified that the LP and HP flare KO drums would be

susceptible to jet fire. However this hazard was considered mitigated against by

the use of deluge and because the system was open to the flare.

Crude cooler - The loss of crude oil cooler inventory was identified and protection

provided using remedial means of isolation.

The remaining M10 areas were no longer mentioned in relation to jet fire.

6.2.2.7 Conclusions

Clearly, jet fire was considered extensively during the design. Indeed the platform is in

some respects designed to resist its effects, i.e. the PFP on the structural members

and the flare boom. Within this is the assumption that the jet fire can last significant

periods of time. Also implicit in the work is the ineffectiveness of deluge on protecting

the affected equipment against jet fire impingement. On the other hand, and of most

relevance to this report, there is no indication that jet fire was ever considered in the

blowdown system design.

In the event the issue seems to have been finally lost when the emergency systems

redundancy was deemed sufficient to perform the intended function, which was to

prevent escalation of a DAE to an RAE through loss of inventory (jet fire is not explicit

but appears to be the cause of the concern). The flaw in the argument is none of the

systems were actually included in the analysis of jet fire escalation or could be shown

to prevent escaltion (although all would be helpful in the situation).

This leaves the platform with the case where a jet fire impinges on a vessel

(particularly the LP separator) with no explicit protection to avoid escalation to at least

a DAE explosion event.

Because of the flaw mentioned above, the FRA did not look at the potential for jet fire

escalation to a RAE so no acceptability criteria were ever established. Potentially

therefore there was a missed RAE event in the analysis.


68

It would be wrong to read this description believing that Hibernia has some design

deficiency. Generally this was the case with all facilities at the time, as the design

codes contained no guidance on how to cope with the jet fire hazard. In fact Hibernia

is much better than most facilities in this regard as will become evident in the following

sections.

6.2.3 Current Requirements of the Design Codes and Practices

6.2.3.1 Mobil (MP 70-P-06, July 1998)

The MEP remains silent on the issue of jet fire and its requirements have not

materially changed since the versions used during design.

6.2.3.2 API 521 (Fourth Edition, March 1997)

The requirements of the API codes has not changed in relation to jet fire since the

design phase and there still is no explicit requirement to design for jet fire events.

API’s position seems to be that jet fire is a low probability event whose effects are

analogous to explosion. These consequences are beyond the ability of a blowdown

system to contain. API were contacted to confirm whether this is the case. They have

responded the issue will be addressed once again during the API 521 revision planned

to commence during 2001. Granherne are pressing API to provide a response within

the timescale of this study.


Once a hazard is identified it does not really matter that the codes of practice are silent

on the requirements to mitigate the hazard and this is the situation the industry finds

itself in with regard to jet fires. We fully expect future versions of API 521 to consider

in more detail the effects of jet fire and we know of at least two other organisations

performing research on the subject (Shell and a Joint Industry Project).

Current industry practice is tending towards incorporating jet fire into the analysis of

new facilities. Granherne know of at least 3 recent projects that were designed with

the assumption of jet fire impingement on equipment was a design criteria.

However, in the absence of prescriptive methods, the way the available research is

used will vary by company although the key aspects are likely to be similar.

Granherne’s approach would follow the following lines:

In terms of methodology the process follows early work by Gayton and Murphy

(Reference 11), who proposed the following methodology:

For each item of equipment define the type of fire (pool, jet, partial engulfment,

total engulfment) likely to affect it.


69

Calculate rate of heat input appropriate to that type of fire.

Calculate the rate of temperature rise of the vessel wall.

Estimate the time to rupture.

If the time to rupture does not meet safety criteria, then design changes may be

necessary to improve the vessel protection.

The sense of the above is self-evident.

Not all vessels will need fire protection. Current studies by Granherne use a more

sophisticated approach than the simple table presented in the QRA, but essentially they

confirm the reasonableness of the early work if only by default. If it can be shown that the

risk of escalation given blowdown is low, then fire protection may be shown not to be

necessary.

Figure 6.3 - Output from a 1-D Heat Up ProgramRESULTS OF HEATUP CALCULATIONS

0.0

100.0

200.0

300.0

400.0

500.0

600.0

700.0

0.0 1.3 2.5 3.8 5.0 6.3 7.5 8.8 10.0 11.3 12.5 13.8 15.0 16.3 17.5 18.8 20.0 21.3 22.5 23.8 25.0 26.3 27.5 28.8 30.0time (min)

he

at

flu

x (

kW

/m2

), t

em

pe

ratu

re l

iqu

id &

va

po

ur

sp

ac

e (

de

g C

)

0.0

20.0

40.0

60.0

80.0

100.0

pre

ss

ure

(b

ara

), s

tre

ss

(%

yie

ld2

0),

yie

ld (

%y

ield

20

)

Heat Input

Liq. Temp

Gas Temp

Press. Bar

% Yield Stress 20C - Applied

% Yield Stress 20C- Strength

The industry is supporting more detailed analysis of fires offshore and performing

experiments to determine how fires behave in confined spaces. These have shown that

fires actually fill the upper space in a module. Also the heating fluxes are considerably

lower than the 300 kW/m2 originally used for analysis. This is because on a platform the

rate of burning is dependent on the ventilation whereas the original figures were from

research in the open air. Although there remains the potential for flame impingement and

engulfment, the implications are that the airflows around the vessel are not so severe,

and deluge systems will still be able to cool the skin. On the other hand, high level

pipework may be more at risk (although pipework is normally considered more robust).


70

One last aspect where change is evolving is in the benefit taken for insulation surrounding

a vessel. As long as it does not catch fire (which is normally the case), and is clad in

steel (rather than aluminium) and the fastening system is secure, the insulation is very

effective at protecting the vessel wall from flame impingement.

Means are now available to calculate the interactions between flame and vessel (and

insulation, if appropriate) in 3D. Whilst confirming the general conclusions of the earlier

modelling, the detail shows effects such as shadowing, liquid level and the lack of heat

removal laterally through the skin.

Figure 6.4 - Temperatures in a Half Filled Vessel Subject to Fire Load - from Heat

Up 3D


71

The results of these calculations feed into the design requirements for the system.

Therefore, current thinking is not so different than in the Hibernia QRA reports in terms

the dependency of the effects of heat on wall thickness, in that thin walled vessels are

likely to fail in less than 15 minutes. Wall thickness is also related to design pressure,

and the flare drums and associated pipework are vulnerable, being low pressure

systems having thin walls. This is why the flare system is specifically mentioned in the

Hibernia safety work. The problem with the early Hibernia work is these effects were

not carried through (in terms of analysis and engineering) to all the areas that were

likely to be affected.

Blowdown, active fire protection and passive fire protection are complementary means

of reducing the risk (loss) from vessel escalation. Primarily this is an asset protection

scheme, since the immediate fatalities in the area will occur before escalation, and

others are likely to be protected in the TSR. The TSR will limit the impact to the

personnel from escalation.

Using this method, the link between the blowdown and other means of protection

would then be explicit in the quantified risk assessment (QRA). This does not appear

to be the case in the Hibernia QRA.

6.2.5 The Effect of Applying Current Best Industry Practise to Hibernia

In this section the effect of applying best industry practise in relation to jet fire is

reviewed. As has been seen, the issue of jet fire causes 3 related aspects to be

considered:

The analysis of jet fire impingement on Hibernia vessels, i.e. what are the

consequences of jet fire impingement on various vessels?

Dependent on the outcome of this will affect the following considerations.

Should anything be done regarding final blowdown pressure and blowdown

duration?

Should the use of other mitigating means be used?

This would have the following effects on Hibernia.

6.2.5.1 Analysis of jet fire impingement on Hibernia vessels


72

There are a number of pressure levels on Hibernia. Within these pressure levels there

is the potential for leak, leak ignition, jet fire and then the potential for escalation.

Usually the worst case scenarios involve the vessels, as these contain the large

system inventories. The worst case scenario is also usually considered to be a leak

from one of the flanges in the pipework around the vessel impinging on the vessel

itself. This means that short jet fire lengths are still severe in effect. One exception to

this is where there is the possibility of a jet fire from a high pressure system impinging

on a low pressure, thin walled vessel. On Hibernia there is such a case which can be

caused by a jet fire somewhere around the HP separator impinging on the LP

separator.

To analyse the situation the following generic case was selected.

The heating of a vessel engulfed by flame was assessed using the package 3-D Heat

Up, a Granherne-developed heat transfer program. The program 3-D Heat Up treats a

source of heat as a flame as a set of discrete emitting “plates” and the receiver also as

a 3-D shape made up of a number of quadrilaterals. The program allows the user to

include details of any insulation on the surface of the vessel. The user can place the

receiving vessel anywhere with respect to the flame, and for the purpose of this study

the vessel was assumed to be engulfed entirely (which would only result from a very

large leak size), as this was the worst case. The diagram below shows a cross section

of the location of the flame.

Table 6.11 Schematic Arrangement of Flame and Vessel

0

10

20

30

40

50

60

70

0 10 20 30 40 50 60 70 80 90 100

(m)

(m)


73

Jet Flame

Vessel

The heat flux within the flame has been found to depend on the wind speed, since

greater wind speeds imply more mixing and this will generate more burning and higher

fluxes. The heat up was therefore assessed for 3 different wind speeds, 2, 5 and 10

m/s, which were calculated to create fluxes of 120, 160 and 180 kW/m2, respectively.

These were thought to be typical of vessels in areas of the platform where ventilation

control of the combustion process was expected. Higher localised fluxes have been

reported in flames in the open (e.g. up to 300 kW/m2), but are not thought to be

applicable in confined spaces. The program has a component that assesses the

residual stresses at elevated temperatures.

The heat response was modelled for the vessels overleaf:


74

Table 6.12 Hibernia Vessel Parameters Summary

Area /

Equipment

Pressure

(kPaA)

Wall thickness

(incl. cladding

where appropriate)

Gas molecular

weight

Insulation

thickness

HP Separator 4240 60.2 21 50mm (steel

supports)

MP Separator 1240 31.6 26 50mm (steel

supports)

LP Separator 210 17.3 48 50mm (steel

supports)

1st Stage

Suction Cooler

210 15.9* 48 Personnel

protection

1st Stage

Suction

Scrubber

160 15.9 36 Insulation height

(High level =

1130 mm)

2nd Stage

Suction Cooler

1125 12.7* 36 Personnel

protection

2nd Stage

Suction

Scrubber


(High level =

780 mm)

3rd Stage

Suction Cooler

3900 41* 22 Personnel

protection

3rd Stage

Suction

Scrubber


(High level =

800 mm)

Injection Stage

Suction Cooler

17120 102* 22 Personnel

protection

Injection Stage

Suction

Scrubber


(High level =

450 mm)

*Cooler head cylinder thickness.

The heat up was then calculated for each of the vessels for two cases, with and

without insulation. Each of the vessels was assumed to be oriented horizontally, but

the results have been checked against vessels oriented vertically (particularly for the

compressor scrubbers). The results of the analysis were charted. The temperatures

do not predict absolutely the potential for vessel failure, since a depressured vessel

will have reduced stresses.

Sample results for an insulated vessel are shown in the figure. These show the

temperatures in the vessel, depicting it as an “unpeeled” skin:


75

Figure 6.5 Sample Output from 3-D Heat Up for Horizontal Vessel Elements

1

4

7

10

13

16 1

9

S1S2S3S4S5S6S7S8S9

S1

0

S11S1

2

S1

3

S1

4

S1

5

S1

6

265

270

275

280

285

290

295

290-295

285-290

280-285

275-280

270-275

265-270

This plot distorts the vessel somewhat, as points at the end, which is dished, are

closer to each other physically than is represented. As can be seen, the temperatures

on the gas side of the vessel are higher than those on the liquid side. This is because

the liquid conducts heat better than the gas, and also because the liquid is a bigger

heat sink. The kink at the far dished end is due to the modelling of conduction through

skin between points that are close together.


76

Bottom, liquid side of vessel

Top, gas side of vessel

Dished end of vessel

Dished end of vesselTemp (K)

The results for each of the three separator vessels, for 3 different heat loads, with and

without insulation are shown below:

Table 6.13 Heat Up Calculation - Separator Peak Vessel Temperatures (K)

Case 180 kW/m2 160 kW/m2 120 kW/m2

60 mm with insulation

(HP Separator)

358 357 355

60 mm no insulation

(HP Separator)

825 786 696


(MP Separator)

372 370 366

32 mm no insulation

(MP Separator)

1063 1024 903


(LP Separator)

397 392 382

17 mm no insulation

(LP Separator)

1287 1269 1078

For the compressor train components runs were performed only at the highest heat flux

(180 kW/m2).

Table 6.14 Heat Up Calculation - Compressor Components Temperatures (K)

Area / Equipment Temperature

without insulation

Temperature

with insulation

1st Stage Suction Cooler 1353 383

1st Stage Suction Scrubber 1313 343

2nd Stage Suction Cooler 1460 449

2nd Stage Suction Scrubber 1039 335

3rd Stage Suction Cooler 1070 448

3rd Stage Suction Scrubber 863 323

Injection Stage Suction Cooler 747 433

Injection Stage Suction Scrubber 628 313

It can be seen from the above that the temperatures of the MP and LP separators and

compressor components up to the 3rd stage suction cooler are well above 500oC if

there is no insulation around the vessels. The highest temperatures are recorded on

the gas side of the vessel, where there is no liquid to act as a heat sink.


77

When the effect of such temperatures on the stresses within the vessels is analysed, it

is found that the vessels are vulnerable to differing degrees. The program identifies

the residual strength left in the vessel at the final temperature, after 15 minutes of

heating. This residual strength takes into account the potential for blowdown of the

vessel to 50% of the design pressure. Thus it is defined as the ratio of the applied

stresses to the vessel at the end of the blowdown period and the residual strength

remaining in vessel at the elevated temperature.

Typical results of the model are shown below:

Figure 6.6 - Sample Result of 3-D Heat Up Calculation for Stresses

(As a percent of Yield Stress at Temperature)1 4 7

10

13

16

19

S1

S100

100

200

300

400

500

600

Str

es

se

s (

% o

f re

ma

inin

g s

tre

ng

th

Vessel Cell No

500-600

400-500

300-400

200-300

100-200

0-100

Once more the greatest effects are seen on the gas side, which is where temperatures

are highest. In the case above failure would occur as the stresses are 5 times the

remaining (residual) strength.

Information can be extracted from the stress results in several formats, such as time to

failure, highest stress and so on. The result format used below shows the ratio of the

applied stress and the percentage of residual strength remaining at time t = 15

minutes.


78

Table 6.15 - Vessel Stress Analysis - Separators

Vessel Stresses at t = 15 minutes as a Percent of Yield Stress at Elevated Temperature

Case 180

kW/m2

160

kW/m2

120

kW/m2

Remarks


(HP Separator)

31 31 31 Vessel remains intact

60 mm no insulation

(HP Separator)



(MP Separator)


32 mm no insulation

(MP Separator)

145 111 52 Vessel can fail at

higher fluxes


(LP Separator)


17 mm no insulation

(LP Separator)

600 511 204 Vessel fails

The results above show that the HP separator does not heat up significantly even

when the effects of the insulation are not included. Consequently the vessel stress as

a percentage of yield stress at the temperature is low. The integrity of the MP and LP

separators, however, is only guaranteed by the insulation. The heat input is much

lower in the insulated case, and given the modest temperature rises in such cases, it

can be said that there will be no threat to the vessels.


79

Table 6.16 Vessel Stress Analysis - Compressor Components

Vessel Stresses at t = 15 minutes as a Percent of Yield Stress at Elevated Temperature

Area / Equipment 180 kW/m2 without

insulation

Remarks

1st Stage Suction Cooler 141 Equipment fails

1st Stage Suction Scrubber 141 Vessel fails

2nd Stage Suction Cooler 720 Equipment fails

2nd Stage Suction Scrubber 145 Vessel fails

3rd Stage Suction Cooler 82 Equipment

remains intact

3rd Stage Suction Scrubber 49 Vessel remains

intact

Injection Stage Suction Cooler <40 Equipment

remains intact

Injection Stage Suction Scrubber <40 Vessel remains

intact

The results show that the lower pressure vessel’s integrity is only ensured by

insulation. The higher pressure equipment, on the other hand, has wall thicknesses

sufficient to survive a jet fire without insulation. By inspection, this suggests that the

current system where the A train injection compressor is blowdown 3 minutes after the

other systems is acceptable (i.e. the vessels should not fail) even if the A train injection

compressor components are engulfed in a jet fire.


80

Also implicit in the ability of the vessel to survive the fire is the necessity for the

staggering system to function as designed. Some concern has been expressed that

the reliability of the system (software, electronics, ESD/PSD and pneumatics) has not

been conclusively demonstrated. Of the failures that could occur, the failure of the

blowdown system to initiate at all is the most serious with the potential, during a fire, to

allow vessel failure and / or escalation through jet fire or explosion. Even outside a fire

situation, the potential for explosion is seriously increased once the contents of the

compressor system begins to vent into the module through the seals as the seal oil

runs out. Of much less concern would be the failure of the staggering system to pause

the A train injection compressor blowdown. In this case the worst event which would

be likely would be abnormally high radiation rates on the platform (dependent on the

wind condition). However, even this benign failure has the potential to escalate if the

initiating cause is an incident involving the LP separator. In this case the coincident

blowdown would add inventory to the area as the back pressure on the LP separator

would be abnormally high. There appears to be good reason, therefore, to perform a

reliability analysis to confirm the system’s ability to function as required.


81

6.2.6 Jet Fire Conclusions

The integrity of the insulation thus is a key issue for the protection of the lower

pressure equipment, namely:

MP separator

LP Separator

1st Stage Suction Cooler

1st Stage Suction Scrubber

2nd Stage Suction Cooler

2nd Stage Suction Scrubber

All of this equipment is insulated in one form or another.

If the insulation remains intact on the vessel under conditions of jet flame engulfment,

and resists the physical impulse from the momentum of the gas jet, then it is likely that

the vessels will not fail. On the other hand, such integrity does not seem to have been

designed into the vessel, and so some upgrading of the protection may be necessary.

The above does not take credit for the presence of deluge. There is still some debate

in the industry on the ability of deluge to mitigate the effects of jet fire, which relate to

how quickly it is applied after the jet fire event has commenced. If the vessel is too

hot, the deluge has difficulty establishing a cooling skin. However, there are a number

of research projects underway which should eventually define the available credit to

take for deluge. For the moment it is sufficient to state that a system with deluge is

much improved over one without.


82

6.3 Blowdown (Depressuring) System Sizing




As mentioned previously, we have been unable to source a copy of the 1990 version

of the above so instead are required to interpret between the 1985 version and the

Draft 1991 version. Using this approach we can estimate the following requirements

at the design stage. Our interpretation of the requirements of the Mobil guide at the

time is:

Vessels shall be depressured to 690 kPag (100 psig) or to 50% of the design

pressure, whichever is smaller. The maximum time allowed to depressure a system is

2 minutes per 3 mm (1/8 in) of vessel wall thickness. Depressuring time of less than

6 minutes need not be used regardless of vessel wall thickness. Depressuring time

shall not exceed 15 minutes, except with Mobil approval.

Vapour depressuring may not be practical when the vessel design pressure is less

than 690 kPag, as piping and valves may become unreasonably large, or when vapour

depressuring load governs the size of the pressure relief and flare headers. When

vapour depressuring is not practical, vessels may be insulated to reduce the vapour

depressuring load or may be protected by other means such as water sprays.

Start pressure for the blowdown was specified as the maximum operating pressure,

which presumably was equivalent to the pressure trip setting.

These latter forms of protection could also be used in lieu of depressuring if designed

according to Mobil guidelines.

Excursions from these requirements required approval from Mobil.


83


During 1991/2 (when these aspects of the design were being finalised) the relevant

version of API 521 was the third edition 1990. This code required the following in

regard to depressuring system sizing.

These systems should have adequate venting capacity to permit reduction of the

vessel stress to a level at which stress rupture is not of immediate concern. For sizing

criteria this generally involves reducing the equipment pressure from initial conditions

to a level equivalent to 50% of the vessel’s design gauge pressure within

approximately 15 minutes. This criterion is based on the vessel wall temperature

versus stress to rupture and applies generally to vessels with wall thicknesses of

approximately 1 inch (25 millimeters) or more. The required percentage depressuring

rate depends on the metallurgy of the vessel, the thickness and initial temperature of

the vessel wall…

Some operating companies limit the application of vapor depressuring to facilities to

facilities that operate at 250 pounds per square inch gauge (1724 kilopascals gauge)

and above, where the equipment and the volume of the contents are significant. Other

companies provide depressuring on all equipment that processes light hydrocarbons,

and they set the depressured level at 100 pounds per square inch gauge (690

kilopascals) or 50% of the design pressure whichever is the lower. The 100 pounds

per square inch gauge (690 kilopascals) level is intended to permit somewhat more

rapid control in which the source of a fire is the leakage of flammable materials from

the equipment to be depressured. On the other hand, in some cases involving

relatively high-pressure vessels that contain relatively large inventories of light

hydrocarbons, depressuring below the 50-percent level within 15 minutes may not be

practical. However, in certain designs this provides an ample margin of safety with

regard to vessel safety from overheating…

API allows the blowdown to commence with the start pressure at initial conditions.


The RABS describes the following project philosophy:

1. Blowdown sections will be depressured from their normal operating pressure to

690 kPag (100 psig) or 50% of the vessels design pressure, whichever is lower.

The maximum time allowed to depressure a vessel/system shall be 2 minutes per

3mm (1/8 in) of vessel wall thickness. Depressuring time of less than 6 minutes

will not be used regardless of vessel wall thickness. Depressuring time shall not

exceed 15 minutes.

(Emphasis added) RABS page 8


84

The emphasis is added to highlight the start pressure considered.

Clearly the Mobil requirements were followed except in the case of the start pressure

for blowdown. Here the RABS suggests the blowdown commences at normal

pressure, whereas the Mobil requirement was from maximum operating pressure.

This issue is confused further by the fact that the calculations for the compressor

system appears to be based on the pressure trip settings suggesting the RABS

contains a typographic error and the design did indeed follow the Mobil guide.

However, the RABS goes on:

2. As a seal oil systems are to be used on the turbine-driven compressors. A more

stringent depressuring design basis then the basis detailed in 1) is required for

blowdown sections which include a compressor.

In order to avoid gas escape along the compressor shaft, compressor sections are

to be depressurised from their initial settle-out conditions to a pressure less than

the static head exerted by the height of the seal oil rundown capacity. This

capacity is defined as the seal oil reservoir volume between the liquid level trip

switch and an empty reservoir. This volume will be sized to allow for an interval of

15 minutes to depressure all of the compressors to 110 kPa (abs).

In order to minimise the peak initial total LP Blowdown flowrate it was agreed that a

staggered compressor blowdown should be used….

The motor-driven gas compressor (K-33-1) is now to utilise a dry gas seal

arrangement instead of a seal oil system. However as the depressuring rate from

this section is low (less than 5% of the peak initial blowdown flowrate) the same

depressuring basis as defined above for the turbine-driven compressor blowdown

sections has been used for this blowdown section.

RABS pages 8 and 9

Due to the selection of the compressor seals, the atmospheric end pressure could not

be avoided unless the system was significantly modified. Otherwise the remaining gas

in the system would spill through the seals into the module when the seal oil ran out.

Of course, more blowdown isolation valves would have avoided the requirement to

depressure the entire compressor system to atmospheric.

The issue of staggering is described in more detail in Section 6.4.

The RABS goes on further:

3. For the production manifolds (HP, MP and test manifolds) and gas injection

manifold it has been agreed that a specific exception to the basis detailed in 1) will

be taken.


85

The reason for this decision is that these sections have very high design pressures

(34400 kPag for the HP, MP, and test manifolds and 45500 kPag for the gas

injection manifold) and contain only pipework. (Note: Manifold pipe wall

thicknesses are in excess of 1 inch.) Therefore, it seems highly conservative to

depressure to 690 kPag in 15 minutes and gives high peak blowdown flowrates

from these sections which form a significant portion of the total HP blowdown

Flowrate.

It was agreed that the application of the philosophy basis identified in 1) was not

intended for this type of high pressure blowdown section. API RP 521 advises that

for high pressure sections including vessels with wall thicknesses of 1 inch or

more, the depressuring basis can be to reduce the pressure from initial conditions

to 50% of the design pressure within approximately 15 minutes. API RP 521 also

notes that on certain high pressure/inventory blowdown sections the depressuring

basis may be reviewed more critically to provide both a practical and safe basis.

The depressuring basis used for the manifolds is as follows:

i) The gas injection manifold will be depressured to 50% of the design

pressure within approximately 15 minutes.

ii) As the initial pressure of the HP, MP and test manifolds is already below

50% of their design pressure they will be depressured to 50% of their

appropriate downstream separators design pressure within approximately

15 minutes.

iii) As the future gas lift manifold wall thickness will probably be less than 1

inch and the current peak blowdown flowrate from this section is not

excessive, this section will still be depressured in accordance with the

basis detailed in 1).

RABS pages 8 to 10

So, virtually every interpretation of API RP 521 that could have been taken made

eventually was. This probably resulted from the ambiguity the recommended practice

contained at the time. Notice also the start pressure of the manifold blowdown which

is again the normal pressure, further highlighting the inconsistency of approach.

Nonetheless, in our view, all of the above interpretations were acceptable as they are

generally conservative. Whether the approach could be described as consistent is

another matter. This issue is described further in Section 6.3.5.

The RABS would also have benefited from a mention of who agreed these issues and

the forum and documentation they were agreed in as the audit trail appears to go dead

after this time.

This result of this approach is shown on the table overleaf.


86

Table 6.17 Blowdown Section Summary

Blowdown

Section No.

Description/Major Equipment Blowdown Start

Pressure (kPa (abs))

Blowdown Final

Pressure (kPa (abs))

Blowdown

Time (mins)

To

Flare

1 Test Manifold #1 4241 2751 15 HP

2 Test Manifold #2 4241 2751 15 HP

3 Well Clean-up Test Manifold 4241 2751 15 HP

4 HP Production Manifolds 4241 2751 15 HP

5 MP Production Manifolds 1241 1101 15 HP

6 Test Separator #1 (D-3105) 4241 791 15 HP

7 Test Separator #2 (D-3107) 4241 791 15 HP

8 Well Clean-up Test Separator

(D-3106)

4241 791 15 HP

9 HP Separator (D-3101) 4241 791 15 HP

10 MP Separator (D-3102) 1241 791 15 HP

11 1st Stage Compressor and

Suction Scrubber

(K-3301 and D-3301)

316 110 15 LP

12 2nd Stage Compressor and

Suction Scrubber A

(K-3302A and D-3302A)

2046 110 15 LP

13 2nd Stage Compressor and

Suction Scrubber B

(K-3302B and D-3302B)

2046 110 15 LP

14 3rd Stage Compressor and

Suction Scrubber A

(K-3303A and D-3303A)

6122 110 15 LP

15 3rd Stage Compressor and

Suction Scrubber B

(K-3303B and D-3303B)

6122 110 15 LP

16 Injection Compressor and

Suction Scrubber A

(K-3304A and D-3304A)

24115 110 12

Note 3

LP

17 Injection Compressor and

Suction Scrubber B

(K-3304B and D-3304B)

24115 110 15 LP

18 Gas Injection Manifold 45601 22851 15 HP

19 Gas Lift Manifold (Note 2) 13801 791 15 HP

20 HP Fuel Gas KO Drum and Fuel

Gas Filter Separators (D-6201,

Z-6201A/B, and Z-6202A/B)

3201 791 12 HP

21 HP Fuel Gas Cooler (E-6201) 3201 791 15 HP

22 Offgas Manifold 4241/1241 791 15 HP

23 LP Fuel Gas KO Drum (D-6202) 621 601 6 LP

24 LP Separator (D3103) 211 Note 1 Note 1 LP

25 Lift Gas Dehydrator and Suction

and Discharge Scrubbers

(C-3801, D-3801, and D-3802)

(Note 2)

17300/13700 791 15 HP

Note 1: The LP Separator operating pressure is already below the level which it should be depressured to. Therefore, only a nominal

blowdown capacity has been taken for this section.

Note 2: The Lift Gas Dehydrator and Gas Lift Manifold are future items.

Note 3: Injection Compressor “A” blowdown is staggered on a 3-minute time delay after other blowdown sections. Therefore, the

blowdown time for this section is only 12 minutes.


87


6.3.3.1 Mobil “Pressure Relief and Vapor Depressuring Systems” MP 70-P-06, July 1998

This Mobil Engineering Practice (MEP) supersedes the earlier engineering guides.

The key requirements concerned with the design of the blowdown system are given

below:

1. Vessels shall be depressured to 690 kPag (100 psig) or to 50% of the design

pressure, whichever is less. The maximum time allowed to depressure a system is

2 minutes per 3 mm (1/8 in) of vessel wall thickness. Depressuring time of less

than 6 minutes need not be used regardless of vessel wall thickness.

Depressuring time shall not exceed 15 minutes, except with Mobil approval.

2. Vapour depressuring may be impractical when the vessel design pressure is less

than 690 kPag (100 psig), because valves and piping may become unreasonably

large and costly. It is also impractical when the vapour depressuring load governs

the size of the pressure relief and flare headers. When vapour depressuring is not

practical, vessels may be insulated (see MP 70-P-05) to reduce the vapour

depressuring load or they may be protected by other means, such as water sprays

(see MP 70-P-01). The use of either of these alternatives requires Mobil approval.

The Mobil practice also gives guidance on blowdown start pressure:

7.2. Depressuring Flowrate

To calculate the vapor flowrate that is needed to accomplish depressuring, the

maximum expected operating pressure of the vessels under consideration shall be

used as the initial pressure and the pressure specified in Section 7.1.1 as the final

pressure.

The practice refers to the API 521 method with regard to depressuring system sizing

for pool fire and the option of controlling the blowdown peak rate by using controlled

blowdown (i.e. reducing the peak rate by control). The remainder of the document in

relation to depressuring is linked to compositional effects that should be considered

during the unsteady state calculations.

Comparisons with the earlier versions of the Mobil practices suggest there has been

no change that would affect the design of the blowdown system.


88

6.3.3.2 API 521 Fourth Edition, March 1997

The relevant part of the Fourth Edition, in relation to the design of the blowdown

system, is given below:

3.19.1 GENERAL

...A vapor depressuring system should have adequate capacity to permit reduction of

the vessel to a level where stress rupture is not of immediate concern. For sizing, this

generally involves reducing the equipment pressure from initial conditions to a level

equivalent to 50% of the vessel design pressure within approximately 15 minutes.

This criteria is based on the vessel wall temperature versus stress to rupture and

applies generally to vessels with wall thicknesses of approximately 1 inch (25mm) or

more. Vessels with thinner walls generally require somewhat greater depressuring

rate...

Where fire is controlling, it may be appropriate to limit the application of vapor

depressuring to facilities that operate at 250 pounds per square inch gauge (1724

kilopascals gauge) and above, where the size of the equipment and volume of the

contents are significant. An alternative is to provide depressuring on all equipment

that processes light hydrocarbons and set the depressured rate to achieve 100 pounds

per square inch gauge (690 kilopascals gauge) or 50% of the vessel design pressure,

whichever is lower, in 15 minutes..." API 521 pages 24 and 25.

There are two issues which derive from the above (and comparison with the earlier

version used during design):

1. The subtle change in the form of words regarding the relevant pressure levels.

As time has passed the requirement of API 521 has hardened and now represents

the clearest idea of the API’s design intent. The practice’s intent can be read as

follows:

To protect against stress rupture:

Systems with design pressure above 1724 kPag should be depressured to

50% of the design pressure.

Systems with design pressure below 1724 kPag need not be depressured.

However, if it is chosen to do so, the final pressure should be 690 kPag or 50%

of the design pressure, whichever is less.

Vessels with wall thicknesses below 1 inch should be considered separately.


89

API were contacted regarding this issue to confirm the above. API were reluctant

to commit to a final response until the meeting of the API 521 Task Force in

October 2000. Granherne are pressing them for a more rapid comment.

2. In this context, what is the meaning of “immediate concern”?

Immediate concern in this instance is 15 minutes. However this hides the fact that

once the stress level in a vessel is reduced by half, the time taken in an API 521

pool fire to heat the (1 inch and over) steel to a temperature at which stress rupture

is likely of the order of hours. The remainder of the blowdown should have been

completed by the time the vessel fails (assuming sufficient fuel to keep the fire

going this long). Based on this scenario the depressuring requirement can be

seen to be very conservative as is evidenced by the fact that some facilities are

allowed to do without depressuring facilities.

Again the start pressure for the blowdown is referred to as initial conditions.


Granherne would apply API requirements as they were intended, i.e:

Systems with design pressure above 1724 kPag should be depressured to

50% of the design pressure (unless there are good reasons otherwise, for

example, the equipment in the HP compression systems which requires a

lower end pressure due to seal oil considerations – see Section 6.3.5 below).

Systems with design pressure below 1724 kPag need not be depressured.

However, if it is chosen to do so, the final pressure should be 690 kPag or 50%

of the design pressure, whichever is less.

Vessels with wall thicknesses below 1 inch should be considered separately.

The above approach would always lead to the minimum sized flare system, indeed it

has the effect of focussing on the most susceptible equipment, thereby applying a high

level of safety.

In passing we should also mention that the API is ambiguous with respect to pipework

engulfed in fire, implying that it does not always need to be depressured. Presumably,

this is because pipework is considered durable in a fire compared to vessels and the

consequences of failure are less than for vessels probably leading to more, fairly

minor, jet fires as the flanges fail. Onshore, this sometimes allows the high pressure

inlet pipework not to be depressured at all (because it is usually located in an area

where fire is unlikely). Offshore this approach is not normally possible and we would

select the depressuring approach from the above. Because of recent accidents where

pipework has catastrophically failed, Granherne expect pipework systems to become

the next target in jet fire analysis and eventually become part of a common

methodology.


90

The start pressure we believe most logical to apply is from initial conditions in cases

where the blowdown is automatically initiated on fire. The probability of coincidence of

high pressure in the system (i.e. just below the PSHH) and fire can be shown to be

very low. The requirement to start at the maximum operating pressure stems from the

era when fire and gas detection systems were unreliable and, as a consequence, did

not initiate automatically. This meant that the fire had the possibility to heat the

system, raising the system’s pressure, prior to manual operator intervention.

Using this approach care would be needed to adjust the calculations should the

pressure profile in the system change significantly.

In passing it should also be noted that some companies adjust the blowdown time

period to remove the stress on all vessels so that they do not rupture in the event of jet

fire impingement. This would normally be required if the vessels were not protected.

This is a particularly expensive way of catering for the jet fire hazard.

Regarding the time to depressure the vessels. The 15 minutes is selected such that

the temperature reached when the stress is halved leaves the vessel not prone to

rupture. Once this satisfactory situation is reached, the vessel continues to

depressure and the period before escalation should lengthen. In other words, if the

initiation of blowdown is timely the period for evacuation should be significantly in

excess of 15 minutes (although it cannot be guaranteed).


Application of the above design practice requirement would significantly reduce the

load on the HP flare as the HP and MP separators could be depressured more slowly.

The target would instead move to ensuring the stress level in the LP separator fell as

quickly as possible as this is the most likely vessel to fail. The existing design case

includes only a nominal depressuring rate (at the time, the clause in the API regarding

thin walled vessels was less than clear). Satisfying the jet fire calculations in Section

6.2.5 should be the target of the revised calculations.

The above would certainly simplify the slightly conflicting requirements of the RABS.

However, it would have very little effect on the ultimate capacity of the system as most

high pressure parts blowdown to the LP flare because of the HP compressors seal oil

system and the requirement to be at atmospheric prior to the oil running out.

Retrofitting dry gas seals to the compressors, or adjusting the sectioning would be the

only ways to significantly increase the latent capacity of the LP flare.


91

Returning to the conflicting requirements in the RABS; it was either acceptable, or it

was not, to blowdown to 50% of the design pressure. If it was (as we believe it was) it

should have been applied to the entire system except where it could be shown

inappropriate, moving the focus to these cases. This was not done, which leaves the

impression of some ‘fitting’ of the requirements to the selected flare boom length.

Nevertheless, this approach, whilst inconsistent, is generally conservative compared to

the intent of the API and therefore a capacity opportunity exists within the system. To

take advantage of the opportunity would require some hardware changes to limit the

blowdown rate to be compatible with the 50% of design end pressure.

As blowdown on Hibernia is initiated automatically there is also a good case for

reducing the severity of the calculated load on the LP flare system by reducing the

start pressure for the compressor blowdown, thus reducing the inventory removed

from the system. This, too, would have the effect of making some spare capacity in

the system. If this were to be implemented no hardware changes would be required;

the new peak rate with the existing system components would be calculated (which

would be less than at the higher pressure), thereby taking up less of the available

system capacity. Without making hardware changes the blowdown would reach the

end pressure in a little less time than 15 minutes.

Prior to agreeing any of the above the permission of Mobil for a deviation to the MEP

would need to be sought.


92

6.4 Compressor Blowdown Stagger




There are no requirements specific to this issue as far as we are aware.

6.4.1.2 Mobil Requirements (EGS 661-1990)




at the design stage. Our interpretation of the requirements of the Mobil guide at the

time is:

The Mobil guide allowed the control of the depressuring rate so as not to exceed the

maximum allowable rate in the flare system, i.e.:

Vapor depressuring valves may restrict the initial depressuring to the capacity of the

closed pressure relief system and flare.

EGS 661-1985

Whilst not explicit this seems to indicate that the use of staggering to achieve this was

acceptable.

6.4.1.3 API (Third Edition, November 1990)

The third edition is silent on the issue of staggering.


Because of the wet seal oil system the compressors were required to depressure to

the LP flare system. The end pressure was required to be atmospheric. This placed a

large load on the LP flare and by 1992 it was realised that LP flare system was unable

to cope with the peak rate (the radiation levels on the platform were too high). The A

train injection compressor stage blowdown was therefore delayed 3 minutes to reduce

the peak rate experienced.

(See Appendix I, calculation 34-006/A Rev 05 for the summary of the loads from the

various areas).


93


The material requirements of either the current Mobil guide or API 521 has not

changed regarding the acceptability of staggering.


The concept of staggering or sequentially depressuring plant has been around the

industry for many years. Whilst not explicitly allowed by the Mobil guide or API

recommended practice nor is it forbidden.

Used most carefully, staggered blowdown is usually reserved for situations where

plant is sufficiently independent that total plant blowdown is not desirable. A good

example of such a situation is a refinery where there are a number of self contained

plant areas, sufficiently independent, and sufficiently far apart that blowdown for a

plant fire in one area would only be desirable in that area.

This is the normal test for the acceptability for staggered blowdown:

The systems should be sufficiently separate such that common mode failure is not

a concern (this would normally require separate PLC control systems and

instrument air supplies).

The systems should be in separate fire areas.

We have seen these requirements in another Operator’s design guidelines the logic

being self-evident.

Generally these requirements defy the application of staggered blowdown to offshore

facilities and in Hibernia’s case neither of these criteria are achieved.


From the above there is clearly concern regarding the staggering of the A injection

compressor blowdown. The case where there is a jet fire around the A train injection

stage (including scrubbers etc.) causing a blowdown of the remaining plant, which is

not on fire is anomalous. In Section 7.2.3 the required blowdown rates required to

avoid stagger are described.


94

6.5 Two-Phase Relief





6.5.1.2 Mobil Requirements (EGS 661-1990)




at the design stage. Our interpretations of the requirements of the Mobil guide at the

time are:

Pressure relief valves shall be sized in accordance with API RP and API STD 526.

Pressure relief valves handling vapour and liquid should be sized according to the two-

phase flowrate from the vessel. Refer to API RP 521 for guidance in determining

vapor and liquid loads from various types of equipment.

EPG 60-B-05 September 1991 page 13

6.5.1.3 API

API RP 14C requires a pressure vessel to have a relief valve sized for full inflow. API

521 requires designers to size the relief valve for closed outlets. The codes are

ambiguous on how such an event will occur. Elsewhere, API 521 (vaguely) refers to

the following:

“The probability of two unrelated failures occurring simultaneously is remote and

normally does not need to be designed for.”

API 521 Third Edition p. 6

To protect a vessel or system from overpressure when all outlets are blocked, the

capacity of the relief device must be at least as great as the capacity of the sources of

pressure. If all outlets are not blocked the capacity of the unblocked outlets may

properly be considered.

API 521 Third Edition p. 8

API 520 also contained various methods for sizing relief valves including a method for

sizing for two-phase relief. The two-phase sizing method relied on calculating the

required orifice required for vapour relief and liquid relief and adding them together.

The API said the following of the method:


95

A reasonable, conservative method of sizing for two-phase liquid/Vapour relief is as

follows:

a) Determine the amount of liquid that flashes by an isenthalpic (adiabatic)

expansion from the relieving condition either to the critical downstream pressure for

the flashed vapour or to the back pressure, whichever is greater.

b) Calculate individually the orifice area required to pass the flashed vapor

component, using Equations 2-7 as appropriate, according to service, type of valve,

and whether the back pressure is greater or less than the critical downstream

pressure.

c) Calculate individually the orifice area required to pass the unflashed liquid

component using Equation 9. The pressure drop (P1-P2) is the inlet relieving pressure

minus the back pressure.

d) Add the individual areas calculated for the vapor and liquid components to obtain

the total orifice area, A, that is required.

e) Select a pressure relief valve that has an effective discharge area equal or greater

than the total calculated orifice area…

API RP 520 Sixth Edition page 37

6.5.2 How the System was Designed for Two-Phase Relief

HP Separator

Combining the requirements described above (and effectively ignoring the full flow

requirement of API RP 14C) gave rise to the following relief valve sizing case.

The HP separator relief valve is dimensioned by the full associated gas rate at design

oil production rate. The most credible scenario that might lead to such a case would

be blockage of the HP separator vapour outlet. This could occur due to maloperation

of an isolation valve or the failure of the pressure control valve in the vapour outlet.

Once this occurred the pressure in the vessel would quickly rise and should cause an

ESD trip. However as is the case with all relief valve sizing cases this trip is assumed

to fail and the relief valve sized for the resulting case.

From a methodological standpoint this is supportable.


96

The two-phase sizing case was selected to consist of a maximum well and an average

well representing the case where two wells fail to shut in (there is a similar case for the

test separator which should be read in this light). Here we partly return to an API

RP14C type case. There is no description in the RABS of how this case might occur

(to us it appears to require the failure of at least 5 ESVs and a high pressure trip). The

calculation method used to find the required area was the additive method contained

in API RP 520 (Fifth Edition). The calculation showed a very much lower required

area than the full associated gas case (3.2 in2 compared to 9.38 in2 for the full

associated gas case).

The HP flare KO vessel was sized to accommodate the resulting liquids from this case

for a 10 minute relief event.

Test Separator

The test separator relief valve follows the above except the sizing case is the two-

phase sizing case.

6.5.3 Current Requirements Of The Codes And Recommended Practices

6.5.3.1 Mobil Requirements (MP 70-P-06)

The Mobil MEP now requires the following:

Pressure relief valves shall be sized in accordance with API RP 520 PT 1 or local

codes, whichever is the more stringent.

MP 70-P-06 page 20

The relevant edition of the API is the Sixth Edition Errata; 1994. This edition retains

the API additive calculation method.

6.5.3.2 API 520 (Seventh Edition, January 2000)

The API has been extensively rewritten with respect to sizing for two-phase

liquid/vapor relief:

3.10.2 A recommended method for sizing pressure relief devices in two-phase

service is presented in Appendix D. The user should be aware that there are currently

no pressure relief devices with certified capacities for two-phase flow since there are

no methods for certification.

Page 55


97

In Appendix D it goes on to state:

D.1.1 The method for two-phase sizing presented in this Appendix is one of

several techniques currently in use and newer methods are continuing to evolve as

time goes on. It is recommended that the particular method to be used for a two-

phase application be fully understood. It should be noted that the methods presented

in this Appendix have not been validated by test, nor is there any recognized

procedure for certifying the capacity of pressure relief valves in two-phase flow

service.

A series of equations based on the Leung omega method are presented. Finally an

alternative method is also mentioned:

D.1.4 A more rigorous approach using vapor/liquid equilibrium (VLE) models

incorporated into a computerized analytical method based on HEM can be considered.

Appendix D page 69


The issue of the new calculation method has caused concern in the industry. The

history of the change stems from some work prepared by DIERS. This group found

that the API method undersized relief valves in two-phase relief cases undergoing

froth reactions. This led to a certain amount of lobbying to have the DIERS method

incorporated in the API. Other groups (presumably aware the DIERS model would not

be appropriate for the oil and gas industry) began to work on models which had the

capacity to predict two-phase relief flows through orifices. These models have been

tested and appear to indicate the API method undersizes orifices in two-phase flow.

Granherne take a pragmatic view of this situation based around the following

arguments:

The earlier API method defines an effective orifice area, which is used to select the

next larger orifice size for installation. The best evidence which recommends this

method is API do not know of a single overpressure failure event to have occurred

since the method was first incorporated into an API code in July 1990 (although the

method has been around much longer than this). Usually a loss event is the

precursor to changing a recommended practice or design code.

Leung omega and HEM methods size the (sharp edged) orifice required to pass

the flow. As valves are not available in all sizes (only the API STD 526

designations) a designer would have to select the next larger size. The valve is

implicitly oversized.

The later Leung omega and HEM work are based around real orifices rather than

the API type of effective orifice area which contains a number of correction factors

which mean an API orifice is bigger than at first sight it seems.


98

Both the new methods are homogeneous and do not account for phase slip (i.e.

the flow through the orifice is sufficiently slow through the orifice to allow the

phases to remain in thermodynamic equilibrium) and therefore must be

approximations. The additive API method is non-homogeneous and has phase

slip inherent in the method, albeit fortuitously.

However, Granherne recognise the valuable research undertaken to date and expect it

to become the basis of relief valve sizing in the future. We expect the sizing method to

adjust as the research is applied to API 526 relief valves when the comparisons will be

much clearer (maybe to the extent that the resulting valve selections are not so

different from the additive method which, for the moment, they are).

None of this, however, protects HMDC (or their advisors) from the difficulties

mentioned in Section 4.1.3 and the requirement to prove the new recommended

practice is not appropriate if it is proposed not to incorporate it. Clearly, it is feasible to

procure a larger relief valve if the calculation check suggests it is necessary. A QRA

will not, in this case, show any improvement in risk profile for the facility. Yet, by

inspection, a valve that is larger than the existing will cope better if the underlying

basis is true and thereby reduce risk in some unquantifiable manner.

Granherne will therefore apply the new requirements to new projects as a matter of

course.


Applying the Leung omega or HEM method has no effect on the sizing selection for

the HP separator. Based on the 40 kbopd original two-phase sizing cases, the full

associated vapour case is still the defining case. The Test separator, however, is a

different matter. It seems that to cater for the new sizing method the valve size should

increase.

The sizing will also depend on the final philosophy selected for the number of wells

which fail to shut-in. HMDC have performed some work on this aspect, including

taking account of the future number of wells and their corresponding rate, which will

need to be incorporated in the updated RABS.

This project can be undertaken when feasible. We also believe that no restrictions to

production need apply in the time it takes to procure the new valves, as it is arguable

that the system is safe by experience.


99

6.6 Design Windspeed and Direction


when Hibernia was Designed







at the design stage.

The 1985 version used a design windspeed equivalent to 93 km/h (57.8 mph or

84.8 fps or 25.8 m/s) if the discharge tip speed was 0.5 Mach. Otherwise MRDC Loss

Prevention Engineering were to be consulted.

By 1991 this was at the point of changing to:

Radiant heat intensities at a design reference point on the platform shall not exceed

the values in Table 2 with the wind in an adverse direction and at maximum

emergency discharge rates. The design wind speed for determining radiant heat

intensities shall be 12.4 km/h (Corrected to 32.2 km/h) (20 mph). The reference point

will be selected, subject to Mobil approval, as the nearest point on the platform that

cannot be readily shielded. Radiant heat intensities shall also be calculated at 67

percent and 133 percent of design wind speed and at other critical points on the

platform to determine what precautions must be taken for flaring during adverse wind

conditions.

Draft EPG 60-B-05 September 1991


The API is silent on the issue of windspeed to use when designing flare stacks/booms.

The only reference anywhere in the document that refers specifically to particular

windspeed is in Appendix C that uses two definitions of windspeed to size a flare

stack.

Design wind velocity is 20 mph (or 29.3 feet per second). (8.9 m/s)

Normal average windspeed is 20 mph (29.3 feet per second) (8.9 m/s)

There is no suggestion (or otherwise) that these figures should be used for flare

system design.


100

6.6.2 The Windspeed Used During Design

The system was designed on the basis contained in the RABS and subsequent

documentation. This is summarised in the following:

a) Windspeed

From the Project Environmental Data Summary (PEDS) the maximum 1 hour

mean wind speed at 140 m above sea level (i.e. at the flare tip location) is

34.2 m/s based on a 1-year return period. The expected frequency shown in

PEDS, is however, less than 0.1% and only quantifiable in directions that would

not adversely affect flare radiation levels.

Therefore, for the purposes of the radiation calculations, a wind speed of 27 m/s

(60 mph), blowing directly towards the platform, will be considered as the worst

case in accordance with the agreed composite specification basis.

RABS page 22

No basis for this figure was given.



The reference to design windspeed is lost in the new document. It may now be linked

to the GKN Birwelco software recommended for the flare sizing task.


The API remains silent on the methodology for selecting design windspeed and retains

the 20 mph level for the calculation examples.


The selection of design windspeed for flare design is usually prescriptively applied by

the Operators. This has arisen probably because there is so little guidance elsewhere

in the national or international codes. A quick survey of projects that Granherne have

been involved with indicates design windspeeds from 10 to 27 m/s (22 to 60 mph) for

offshore locations, none of which appear to have been set by using a constant

methodology.


101

However, if Granherne were to choose a design windspeed it would be selected on a

probabilistic basis. In a fairly arbitrary sense (as we have not quantified the effect on a

number of facilities of using this method), we too would choose a windspeed based on

the maximum 1 hour mean wind speed at the flare tip location with a probability of

0.1% based on a 1-year return period (i.e. in a similar way to the RABS). The reason

for selecting this yearly approach is because it seems reasonable. However, we

would not necessarily limit the direction to directly onto the platform if a slight deviation

produced a significantly higher windspeed. This would be the absolute maximum we

would consider. If the resultant windspeed were higher than 27 m/s (60 mph) we

would limit consideration to this level and the flare boom would be dimensioned on this

case. This windspeed is derived from Granherne experience of the usual limit placed

on helideck operations. Beyond this level it is no longer safe to be on deck (see also

Section 8.0).

In saying the above, Granherne would also have sympathy for any situation where a

less onerous windspeed were selected; the codes could be interpreted to allow it.


In the event, Granherne’s current best practice would have had a very minor effect on

the flare boom length or (as is the case now the platform is constructed) the design

rates allowable.

Review of the Project Environmental Specification enables the following comparison to

be made:

Table 6.18 Comparison of Design Windspeed Criteria with Granherne Best

Practice

RABS (m/s) Granherne (m/s)

Most Adverse Direction (directly

onto the platform)

Not identified 24.2

Most Adverse Windspeed in an

on platform direction

34.2 from NW (not analysed) 34.2 from NW (analysed and not

as extreme as above)

Design figure 27 24.2

The windspeeds are very similar to the RABS criteria. The effect of using these

windspeeds is demonstrated in Section 8.0.


102

6.7 Acceptable Flare Radiation Levels

6.7.1.1 Canada Oil and Gas Installations Regulations (January 1991, Draft)

Recapping the requirements:

(9) Every gas release system shall be designed and installed so that, taking into

account the prevailing wind conditions, the maximum radiation on areas where

personnel may be located, from the automatically ignited flame of a flare or vent,

will be

(a) 6.3 kW/m2, where the period of exposure will not be greater than one minute;

(b) 4.72 kW/m2, where the period of exposure will be greater than one minute but

not greater than one hour; and

(c) 1.9 kW/m2, where the period of exposure will be greater than one hour.





at the design stage.

The 1985 version used the criteria given in Table 6.10 in the calculation of flare stack

height.

Table 6.19 Allowable Radiant Heat Intensities Excluding Solar Radiation (1985)

Heat Intensities

Allowed, K

W/m2

Location

1580 Areas where personnel must remain at their posts

2365 Storage tanks containing volatile material, and control rooms

4730 Areas where escape of personnel is possible in several minutes

6300 Open areas where refinery personnel can be exposed up to one minute

with appropriate clothing

9465 Areas where protection or shielding from the radiant heat is available to

refinery personnel in six seconds or less (except for control rooms or for

non-combustible equipment and facilities)


103

By 1991 this was at the point of changing to:

The flare stack location shall be determined by allowable radiant heat intensities at

various critical points on offshore platforms or processing facilities. It shall be

calculated in accordance with API RP 521 and as modified by this guide.

The modifications to API RP 521 the 1991 guide refers to are given in Table 6.11

below.

Table 6.20 Allowable Radiant Heat Intensities (1991)

Heat Intensities

Allowed, K

W/m2

Condition(1)

1580 For continuous flaring operations in areas where personnel must remain at

their work stations without shielding but with appropriate clothing

1580 Emergency flaring for several minutes(2) - personnel without appropriate

clothing

790 Continuous flaring(2) - personnel expected to wear appropriate clothing

3155 Emergency flaring up to one minute(2) - personnel without appropriate

clothing

Notes

1. In areas where personnel can be exposed to higher radiation intensities, heat shielding must be

provided and also for equipment and structure as necessary.

2. In areas where personnel are not expected to wear appropriate clothing (i.e. coveralls, boots,

gloves, hard hats) allowable radiation levels have been reduced by a factor of two.


104

API RP 521 (Third Edition, November 1990) recommends the following:

Table 6.21 Recommended Design Flare Radiation Levels Excluding Solar

Radiation (API RP 521)

Permissible Design

Level (K)

KW/m2

Location

15.77 Heat intensity on structures and in areas where operators are not likely to be

performing duties and where shelter from radiant heat is available (for example,

behind equipment).

9.46 Value of K at design flare release to any location to which people have access (for

example, at grade below the flare or a service platform of a nearby tower);

exposure should be limited to a few seconds, sufficient for escape only.

6.31 Heat intensity in areas where emergency actions lasting up to 1 minute may be

required by personnel without shielding but with appropriate clothing.

4.73 Heat intensity in areas where emergency actions lasting several minutes may be


1.58 Value of K at design flare release to any location where personnel are continuously

exposed.

6.7.2 The Radiation Levels Used in the Design

The design used the following radiation levels, derived from Draft Canadian

Legislation, as outlined by the RABS:

Table 6.22 Radiation Flux Limits Excluding Solar Radiation

Permissible Design

Level (K)

KW/m2

Conditions

6.3 Heat intensity in areas where emergency actions lasting up to 1 minute may be


4.72 Heat intensity in areas where emergency actions lasting up to several minutes may

be required by personnel without shielding but with appropriate clothing.

1.9 Value of allowable radiation level at design flare release at any location where

personnel are continuously exposed, i.e. helideck


105

In addition the following statement is included in the RABS:

In addition to the above radiation limitations HTPT advised that the maximum radiation

level experienced on the platform escape routes is not to exceed 1000 Btu/ft2 h (3.16

W/m2) for periods over 1 minute of exposure.

RABS page 22

The design calculations for the worst emergency flaring case (total platform blowdown)

and a flare boom length of 115m resulted in the following radiation levels:

Approximately 3.16 kW/m2 at the north side M10 weather deck

Approximately 6.30 kW/m2 at the drilling derrick crown block

Approximately 4.72 kW/m2 at the drilling derrick finger board



The new document has the following recommendations on thermal radiation levels.

Table 6.23 Allowable Radiant Heat Intensities in W/m2 Excluding Solar Radiation

Appropriate Clothing* Without Appropriate Clothing*

Continuous Release 1105 790

Emergency Releases

Travel time to Shelter

6 Sec 9465 3150

1 Min 6300 3150

3 Min 4730 1580

No Shelter Available 1580 790

Equipment Only

Exposure15770

Volatile Liquids Tanks,

API Separators, CCB

2365


106


The recommendations on flare thermal radiation levels in the new addition of API

RP521 remain the same as the previous version.


We have mentioned earlier in this report that best practice is subjective to some

extent. Where issues are not subjective are in matters of law. Once a requirement

passes into law, as have the Canadian regulations, by meeting those requirements, an

owner has effectively discharged their responsibilities. As is also customary in matters

of precedence, national regulations always supercede recommended practices.

Normally there is actually little difference between the two requirements and this is

where we find ourselves in the Hibernia context. This is demonstrated in the following

table:


107

Table 6.24 Summary of Flare Radiation Requirements for Hibernia

Canadian Regulations

Impairment Criteria

Equivalent API RP521

Remarks

Maximum radiation on areas where the period of exposure will be greater than one hour

1.9 N/A 1.58 Apply Canadian regulations.

Helideck operable for at least 2 hours. Inoperability may result from…thermal radiation over 3.2 kW/m2

(Impairment Criterion 4)

Silent 3.2 Not specifically mentioned

Not used as a normal radiation level.

Maximum radiation on areas where the period of exposure will be greater than one minute but not greater than one hour

4.72 N/A 4.73 Apply Canadian regulations.

Maximum radiation on areas where the period of exposure will not be greater than one minute

And,

Escape routes from all parts of the platform to the TSR… to remain passable for 30 minutes…An escape route may be made impassable by:

Thermal radiation over 6.3 kW/m2

if unprotected:


6.3 6.3 6.3 Apply Canadian regulations.

Maximum radiation on areas where the period of exposure will not be greater than a few seconds

(In this case the area in question is normally accessible)

Silent N/A 9.5 Apply API requirements in absence of Canadian regulation.

The actual wording of API 521 is:

Value of K at design flare release to any location to which people have access (for example, at grade below the flare or a service platform of a nearby tower); exposure should be limited to a few seconds, sufficient for escape only. (Note 1)

Escape routes from all parts of the platform to the TSR… to remain passable for 30 minutes…An escape route may be made impassable by:

Thermal radiation over 12.5 kW/m2 to the outside of the escape route if protected by cladding:


Silent 12.5 Not specifically mentioned

Not used as a normal radiation level.

Maximum radiation on areas where shelter is present.

(In this case the area in question is not normally accessible)

Silent N/A 15.8 Not used as a normal radiation level.

The actual wording of API 521 is:

Heat intensity on structures and in areas where operators are not likely to be performing duties and where shelter from radiant heat is available (for example, behind equipment). (Note 1)

Notes:

1_) On towers and elevated structures where rapid escape is not possible, ladders must be provided on

the side away from the flare, so the structure can provide some shielding when K is greater than …

6.3 kilowatts per square meter.


108

The allowable radiation levels on Hibernia will have to be selected from within these

requirements.


The application of Canadian regulations and interpretation of the guidelines in API

RP521 where the Canadian requirements are silent (replicated in Table 6.24) for the

Hibernia platform gives the following allowable thermal radiation levels at various parts

of the platform:

Crown Block 9460 W/m2

The crown block falls into the category an area to personnel have access (i.e. a

service platform of a nearby tower); where exposure can be limited to a few seconds,

sufficient for escape only.

It could even be argued that a more extreme limit at this point could be used: The

Damage / Impairment criterion No. 3 indicates that a value of 12500 W/m2 may be

appropriate for the area under consideration. The criterion is specifically aimed at

escape routes protected by cladding but could equally be applied to the drilling derrick

which is partially enclosed and offers any operator working in the area the opportunity

to shelter behind a clad structure for the duration of the emergency. A reference for

the figure of 12500 W/m2 cannot be found in the guides and practices referenced

above although a somewhat worse value of 15800 W/m2 can be found in the API

which is allowed only in an area where shielding exists. These requirements are

included for information only.

In the original design it appears an unnecessarily conservative approach, which did

not recognise the presence of shielding, was applied to this area which limited the

radiation level to 6300 kW/m2.

Weather Deck 6300 W/m2

The weather deck and monkey board falls into the category of an area where

emergency actions lasting up to one minute may be required by personnel without

shielding but with appropriate clothing. It is expected that personnel on the weather

deck would be appropriately clothed and in the event of an emergency blowdown

would be able to leave either leave the deck in a minute or less or alternatively find

shelter in the same time period.

In the original design, values of 3200 and 4720 kW/m2 respectively were applied to

these areas. The former resulted from the HTPT note attached to the table which

outlined the explicit design requirements and, from the above, is a radiation level

allowable for 2 hours in an emergency. The latter neither recognises the escape

ability from the monkey board nor the shielding. Both cases therefore appear

unnecessarily conservative.


109

Helideck 3200 W/m2

The helideck does not really fall into any specific category as defined in the guides and

practices reference above but it could be argued that it loosely falls into the category of

an area where personnel are continuously exposed during an emergency (for up to

two hours) and therefore the value of 3200 W/m2 is chosen. This corresponds with the

Damage / Impairment criterion No. 4 which indicates that a value of 3200 W/m2 is

appropriate for the helideck which is based on Canadian regulations.

Weather Deck (Continuous flaring) 1900 W/m2

For a true (non-emergency) continuous release, e.g. production flaring, a figure of

1900 kW/m2 should be used (again in accordance with Canadian regulations).

Of the above, the most important radiation level is likely to be the continuous flaring

case as it will be the most persistent (occasionally). The other radiation levels are only

approached during a platform blowdown and therefore are short duration (only

seconds) and will only be felt if a coincident severe adverse wind occurs during the

event.

Using this radiation level and location as the design case ensures that the helideck will

experience very much lower radiant rates during continuous flaring.

The radiation levels used in the flare operating envelope calculations, described fully in

Section 8.0 below, have used the above thermal radiation limits to determine the

allowable maximum flaring capacity for the ‘As Built’ flare for two windspeeds.

The results of these calculations are discussed in detail in Section 8.0 but the principle

conclusion is that if the best practice radiation levels are applied as defined above then

the flare system capacity would be approximately 200% of the current design load in

terms of thermal radiation only. The effect on hydraulics in the system for this capacity

increase has not been studied at this time.

This section should be read in conjunction with Section 8.4.


110

6.8 Challenge Issues Resulting from the Technical Audit of the Design Calculations

Issue 34-005/2 - Jet fire scenario was not taken into account for in the design of

the blowdown system.

This issue is addressed in Section 6.2.

Issue 34-005/4 - Were fire areas used for total blowdown rate?

A simple approach to blowdown was used where the entire all equipment to be

depressured was assumed to be on fire. This is equivalent to the entire M10 module

being on fire, something which is very unlikely and if it occurs will be catastrophic.

More conventionally the platform is separated into fire areas. In this case the

blowdown valves are sized to cater for the fire case. However, the combined case is

not normally the sum of all the areas on fire and some effort is instead focussed at the

selection of a realistic worst case. The worst case is represented by the fire occurring

in the area which adds most to platform load coincident with the resultant rates from

the blowdown valves for the non-fire areas are added. These latter rates are less than

the rate that would be experienced in a fire case and the overall blowdown load is

more accurately represented. In this case we have been unable to locate fire area

drawings which forces the M10 fire case to remain the design case.

Issue 34-006/2 - Is correct isentropic efficiency used?

The isentropic efficiency specified when performing blowdown simulations affects both

the downstream blowdown temperature of gas and equipment but also the upstream

vessel wall temperature. An isentropic efficiency of 1 simulates perfectly isentropic

expansion of the gas and gives the worst case (i.e. lowest) temperatures. An

isentropic efficiency of 0 simulates perfectly isenthalpic expansion of the gas and gives

the best case (i.e. highest) temperatures. For blowdown of a vessel or system where

the feed to the vessel has been stopped the expansion of the gas is somewhere

between isentropic and isenthalpic. The selection of the isentropic efficiency is usually

based on project philosophy and experience.

The blowdown simulations performed for these calculations used an efficiency of 0.5.

There is no indication in the calculations or simulation outputs for the basis of this

selection. The selection of an efficiency of 1 is unrealistic, but a more usual figure to

use is 0.7 minimum which would lead to lower blowdown temperatures.

The impact of lower blowdown temperatures is twofold. The first concerns the

materials of construction of the flare system itself. The flare system appears from the

‘As Built’ P&IDs to constructed of LTCS with a minimum design temperature of -45oC.

The second impact is in areas of the process where hydrates can form.


111

From the calculations reviewed problems of both hydrate formation and flare design

temperature only occur if blowdown is initiated after a delay with of the plant

maintained at pressure during the upset. Calculation 34-010 / A and 34-060 / B

address this problem but the calculations do not give any specific conclusions on the

allowable delay. Calculation 060 / B concludes that for the settleout pressures used

there is a huge spread of allowable delay periods depending upon environmental

conditions and whether insulation is installed. Current platform design philosophy is to

depressurise after 1-2 hours. If lower blowdown temperatures are expected then this

philosophy may have to be reviewed. See Issue 34-010/1 for further details.

Precautions against hydrate formation can be taken and these are discussed further

below.

For more discussion regarding this see the related discussion in Section 5.2.1.2 item

34.010/1.

Issue 34-006/3 - Is design case too extreme?

This concern relates to the high start pressure used for the blowdown calculations.

See Section 6.3 for the recommended solution.

Issue 34-042/2 - Validity of staggering blowdown. Were the systems sufficiently

independent?

This aspect is covered in Section 6.4.

6.9 Miscellaneous Issues

6.9.1 Insulation

The presence of satisfactory insulation on the vessels will allow substantially reduced

depressuring rates compared to those used during design. This aspect should be

checked in detail when the insulation on the vessels is reviewed.


112

7.0 AS-BUILDING THE FLARE SYSTEM

7.1 Introduction

The Hibernia platform was built incorporating features for future equipment; this

included future capacity built into the flare system. As some of these projects are no

longer foreseen this section looks to remove their effect from the currently installed

flaring cases. This, in effect, will result in a system whose design cases are “as-built”.

The difference between the design capacity and the “as-built” capacity is the capacity

available for future projects, including those that were originally foreseen.

7.2 As-built and Design Capacity

The following table summarises the initial relieving capacity by area considered during

the design phase as well as the results of removing the requirements for future

equipment and potentially the 3 minute stagger on the injection compressor ‘A’

blowdown.

Table 7.25 As Built and Design Capacity

Case Scenario LP Flare Load

kg/h

HP Flare Load

kg/h

1 Design Blowdown Rate as per Relief &

Blowdown Study Report Rev C1

89,601 133,616

2 ‘As Built’ i.e. As Case 1 with Future Equipment

Removed

89,601 94,843

3 As Case 2 with 3 min Stagger Removed 126,291 94,843

The table below summarises the effect on thermal radiation impingement at various

points on the platform for the cases described in Table 7.1 with the original design

wind speed of 27 m/s blowing in a northerly direction.

Table 7.2 Thermal Radiation Impingement on Platform Areas

Case Crown Block

W/m2

Weather Deck

W/m2

Helideck

W/m2

1 6152 2911 1034

2 5623 2700 860

3 8838 3146 1314


113

The capacity of the flare system is essentially decided by the allowable thermal

radiation impingement on the platform. The different levels of thermal radiation and

their limitations on working and escape routes are discussed in Section 6.7. Based on

the original design radiation levels:

Crown Block 6300 W/m2


Helideck 1900 W/m2

Then the following deductions can be made from the resulting thermal radiation

impingement on the platform for the 3 flare operating cases described in Table 7.1.

7.2.1 Case 1 - Design Blowdown Rate (as RABS Rev C1)

The results for this case indicate that the thermal radiation impingement at the crown

block is close to the limiting value of 6300 W/m2. This is what we would expect as the

flare stack lengths was effectively sized for this flaring scenario. The thermal radiation

impingement at the weather deck and helideck are within the limits stated above

7.2.2 Case 2 - ‘As-Built’ - i.e. As Case 1 with Future Equipment Removed

The results for this case indicate, as expected, that removing the load assigned for

‘future’ equipment from the HP flare gives some margin for increased flare load

generated by future projects / platform modifications. The margin is available for

thermal radiation at all platform areas discussed but, by inspection, the limit is

expected to occur at the crown block. It should be noted that any projects which

generate extra coincident LP blowdown load on the flare will have a greater effect on

thermal radiation impingement than that for HP blowdown due to the nature of the flare

systems. Therefore capacity, in terms of mass flowrate, liberated from the HP flare is

not necessarily available in full for the LP flare system.

7.2.3 Case 3 - As Case 2 with 3 min Stagger Removed

This case investigates the effect on the ‘As Built’ flare (i.e. with loads from ‘future’

equipment removed) of removing the 3 minute stagger between blowdown initialisation

and the blowdown of the Injection Compressor ‘A’ system. This is potentially a

modification that HMDC would consider making in the future. It can be seen from the

results, however, that though the thermal radiation levels at the weather deck and the

helideck are acceptable, the level at the crown block is greater than the current design

limit based on the criteria above. This would likely be acceptable if the shielding

around the drilling derrick structure was taken into account.


114

8.0 RISK MANAGEMENT IN RELATION TO FLARING EVENTS AND

WIND CONDITION

8.1 Introduction

In Section 6.6 the issue of design windspeed was discussed. In this section we look at

the various windspeeds to show the effect it has on the radiation envelope on

Hibernia.

The selected windspeeds are:

The most adverse windspeed (with a probability of 0.1% and a return period of 1

year) in a direction which could influence the radiation levels on the platform. In

this case the windspeed = 34.2 m/s from the NW.

The most adverse windspeed (with a probability of 0.1% and a return period of 1

year) directly onto the platform. In this case the windspeed = 24.2 m/s from the N.

The original design windspeed from the RABS. In this case the windspeed =

27 m/s.

The results are shown below:

8.2 Potential Flare Envelope based on Total Blowdown Scenarios

8.2.1 Determination of Blowdown Load Basis

A basis for constructing an operating envelope for this study had to be developed with

no specific modification projects in mind. The task is complicated by the fact that there

are two flare systems, the LP flare and the HP flare. An increase in flare load has a

different consequence depending upon the particular flare affected, due to the different

nature of the flares. The HP flare utilises a sonic tip and therefore has a flame that

burns much more efficiently and is stiffer than the flame developed by the pipe flare tip

on the LP network.

Given the above, the only sensible approach to preparing an operating envelope was

to base it on multiples of the radiation case defining case, i.e. total platform blowdown

giving coincident LP and HP flare release as defined in the Relief and Blowdown Study

Report. The relieving loads on each flare for the design flare capacity case are given

as Case 1 in Table 7.1 above.

Flaresim calculations were performed for flare loads ranging from 50% to 500% of the

design case to enable an envelope to be defined. In reality, the systems would be

hydraulically limited long before these higher rates were achieved. The relief rate for

each flare was maintained in the same proportion as the design case and fluid

properties also remained constant for all capacities considered.


115

8.2.2 Determination of Continuous Load Basis

The Relief and Blowdown Study Report identified that the worst case continuous

flaring occurs when there is a relief load on both the HP and LP flare systems. This

occurs either at start up or when the compression train is lost for any reason.

To determine the continuous relief load, we considered the loss of HP and LP

compression at 100% production causing the associated gas from the to spill-off to the

HP and LP flares, from the HP and MP separators and LP separator respectively. To

generate the input data we used a simulation taken from the recent debottlenecking

project (Case 3, a case which included Avalon production) with a total production of

200 kbopd (taken in this instance as 100% capacity) and used this to calculate spill-off

rates to each flare system if the compression train is disconnected. At 100%

production, therefore, the flare loads are 359.2 Te/h (MW 21.3) to the HP flare and

50.4 Te/h (MW 45.0) to the LP flare. Note that in this case the HP flare rate exceeds

current capacity.

Flaresim calculations were performed for flare loads ranging from 30% to 100% of the

above determined rates to enable an envelope to be defined. In reality, the systems

would be hydraulically limited before these higher rates were achieved. The relief rate

for each flare was maintained in the same proportion as the 100% case and fluid

properties also remained constant for all capacities considered.

8.2.3 Determination of Allowable Thermal Radiation Impingement on the Platform

The flare capacity envelope for any area on the platform is very much dependent on

the maximum allowable thermal radiation impingement at that particular area. For this

study the following limiting thermal radiation impingement levels were used to

generate the operating envelopes:

Crown Block 9500 / 12500 W/m2


Helideck 3200 W/m2

Continuous *1900 W/m2

* API suggests a lower figure (1580 kW/m2) is appropriate. In this case, because of Canadian regulations,

this is ignored.

For the areas with two thermal radiation levels given, the lower figure refers to the

‘Best Practice’ value as identified in Section 6.7.5 above and the upper figure is the

allowable thermal radiation level in accordance with the Impairment / Damage criteria

which is included for information. The reasoning behind the choice of limiting thermal

radiation levels and their effect on personnel and structures are discussed in detail in

section 6.7 above.


116

8.2.4 Other Calculation Criteria

Relief gas compositions for the total platform blowdown design case were taken from

the Kaldair Design Data Dossier (Reference 13) and used to generate the fluid

properties used in the Flaresim simulations.

Windspeeds used for the calculations are 27 m/s (original design) and 24.5 m/s

(determined from environmental data). Both these wind are blowing in a Northerly

direction, i.e. directly back onto the platform from the direction of the flare boom. The

reasoning behind the choice of windspeeds is discussed in detail in section 6.6 above.

Also considered was the 34.2 m/s NW wind considered in the RABS. The results of

the analysis indicated very similar results as the 24.5 m/s windspeed.

8.3 Results


The results of the study are presented in graphical form in figures 8.1 to 8.10 below.

The curves on each graph represent the distance of the isopleth from the flare tip

varies with blowdown rate. The isopleth under consideration depends upon the area

of the platform under consideration as described above. The distance of isopleth to

flare tip is measured in the direction of the platform area under consideration. For

example, Figure 8.1 shows the distance of the 12500 W/m2 from the flare tip in the

direction of the crown block. The Horizontal line represents the actual distance of

crown block from the flare tip. Where the two intersect gives the blowdown rate, as a

percentage of design, which would result in thermal radiation of 12500 W/m2 impinging

on the crown block.

For information, although not applicable to emergency relief, the radiation distance to

the 1900 kW/m2 level is also given. This would be the type of figure that would be

considered allowable around the TSR.


117


118

Figure 8.1 Envelope of Operability for Crown Block with Limiting Thermal Radiation 12500 W/m2 with Northerly Wind 24.5 m/s

0.0

20.0

40.0

60.0

80.0

100.0

120.0

140.0

0% 50% 100% 150% 200% 250% 300% 350% 400% 450% 500% 550% 600%

% Design Case Blowdown Rate

Dis

tan

ce f

rom

Fla

re T

ip

Crow n Block Minimum Distance from Tip

12500 W/m2 Isopleth Distance from Tip


0.0

20.0

40.0

60.0

80.0

100.0

120.0

140.0

160.0

0% 50% 100% 150% 200% 250% 300% 350% 400% 450% 500% 550% 600%


Dis

tan

ce f

rom

Fla

re T

ip




119

Figure 8.3 Envelope of Operability for Weather Deck with Limiting Thermal Radiation 6300 W/m2 with Northerly Wind 24.5 m/s

0.0

20.0

40.0

60.0

80.0

100.0

120.0

0% 50% 100% 150% 200% 250% 300% 350% 400% 450% 500% 550% 600%


Dis

tan

ce f

rom

Fla

re T

ip


Weather Deck Minimum Distance from Tip

Figure 8.4 Envelope of Operability for Helideck with Limiting Thermal Radiation 3200 W/m2 with Northerly Wind 24.5 m/s

0.0

20.0

40.0

60.0

80.0

100.0

120.0

140.0

160.0

180.0

200.0

0% 50% 100% 150% 200% 250% 300% 350% 400% 450% 500% 550% 600%


Dis

tan

ce f

rom

Fla

re T

ip

3200 W/m2 Isopleth Distance from TipHelideck Minimum Distance from Tip


120


0.0

50.0

100.0

150.0

200.0

250.0

0% 50% 100% 150% 200% 250% 300% 350% 400% 450% 500% 550% 600%


Dis

tan

ce f

rom

Fla

re T

ip


Helideck Minimum Distance from Tip


0.0

20.0

40.0

60.0

80.0

100.0

120.0

140.0

0% 50% 100% 150% 200% 250% 300% 350% 400% 450% 500% 550% 600%


Dis

tan

ce f

rom

Fla

re T

ip




121

Figure 8.8 Envelope of Operability for Weather Deck with Limiting Thermal Radiation 6300 W/m2 with Northerly Wind 27.0 m/s

0.0

20.0

40.0

60.0

80.0

100.0

120.0

0% 50% 100% 150% 200% 250% 300% 350% 400% 450% 500% 550% 600%


Dis

tan

ce f

rom

Fla

re T

ip




0.0

20.0

40.0

60.0

80.0

100.0

120.0

140.0

160.0

0% 50% 100% 150% 200% 250% 300% 350% 400% 450% 500% 550% 600%


Dis

tan

ce f

rom

Fla

re T

ip




122


0.0

20.0

40.0

60.0

80.0

100.0

120.0

140.0

160.0

180.0

200.0

0% 50% 100% 150% 200% 250% 300% 350% 400% 450% 500% 550% 600%


Dis

tan

ce f

rom

Fla

re T

ip




0.0

50.0

100.0

150.0

200.0

250.0

0% 50% 100% 150% 200% 250% 300% 350% 400% 450% 500% 550% 600%


Dis

tan

ce f

rom

Fla

re T

ip




The results of the continuous flare relief study are presented in graphical form in

figures 8.11 and 8.12 below. The curves on each graph represent the distance of the

1900 W/m2 isopleth from the flare tip varying with continuous relief rate. The distance

of isopleth to flare tip is measured in the direction of the platform area under

consideration, in this case only the weather deck is considered. The horizontal line

represents the actual distance of weather deck from the flare tip. Where the two

intersect gives the continuous relief rate, as a percentage of design, which would

result in thermal radiation of 1900 W/m2 impinging on the weather deck.

The figures in parentheses on the x axis represent the combined LP and HP flare

mass flowrate under consideration e.g. the mass of gas flared if the platform is

operating at 100% capacity (considered to be 200 kbopd) is 409.6 Te/h with a pseudo

molecular weight of 22.8. Note that the mass flowrates stated become invalid if the

ratio of HP to LP flare load differs from that considered here (see Section 8.2.2 above

for further details).


123


124

Figure 8.11 Envelope of Operability for Weather Deck with Limiting Thermal Radiation 1900 W/m2 (Continuous)

with Northerly Wind 24.5 m/s

0.0

20.0

40.0

60.0

80.0

100.0

120.0

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100% 110%

% Design Case Continuous Flaring Rate (Total Flare Mass Rate, Te/h)

Dis

tan

ce f

rom

Fla

re T

ip, m



Figure 8.12 Envelope of Operability for Weather Deck with Limiting Thermal Radiation 1900 W/m2 (Continuous)

with Northerly Wind 27.0 m/s

0.0

20.0

40.0

60.0

80.0

100.0

120.0

0% 10% 20% 30% 40% 50% 60% 70% 80% 90% 100% 110%

% Design Case Continuous Flaring Rate (Total Flare Mass Rate, Te/h)

Dis

tan

ce f

rom

Fla

re T

ip, m



(122.9) (163.8) (204.8) (245.8) (286.7) (327.7) (368.7) (409.6)(81.9)(41.0)(0) (450.6)

(122.9) (163.8) (204.8) (245.8) (286.7) (327.7) (368.7) (409.6)(81.9)(41.0)(0) (450.6)

8.4 Flare Envelope Conclusions

8.4.1 General

An immediate and obvious conclusion which can be drawn from the results of this

study is that a wind speed of 27 m/s gives higher thermal radiation levels on the

platform than the 24.5 m/s wind speed for the relief cases considered here.

However if we analyse the results given for the wind speed of 24.5 m/s, the Granherne

best practice wind speed as defined in Section 6.6, figures 8.1 to 8.5 and 8.11 above,

the following conclusions can be drawn for each flaring scenario:


Maximum allowable thermal radiation at the crown block - 9500 W/m2

For this case the thermal radiation on the crown block is the limiting factor. An initial

blowdown rate of approximately 150% of the design rate can be tolerated before the

limit of 9500 W/m2 is limit is exceeded in this area.

The other cases considered are less onerous, with the 3200 W/m2 isopleth impinging

on the Helideck at around 350% of the design blowdown rate and the 6300 W/m2

isopleth impinging on the weather deck at around 370% of the design blowdown rate.

The above suggests considerable capacity is inherent in the system dependent on the

final basis selected. However, caution should be exercised as this apparent capacity

will change dependent on the detail of the project which actually utilises the apparent

capacity. In other words the absolute capacity will only be confirmed once the LP and

HP rates are fully defined and detailed calculations are performed for the modification

under consideration.


For this case the maximum allowable thermal radiation of 1900 W/m2 at the weather

deck is considered to be the limiting factor. For a northerly wind blowing at 24.5 m/s, a

platform production rate of 62% of the design rate (considered to be 200 kbopd) can

be tolerated before the limit of 1900 W/m2 is limit is exceeded in this area.


125

Here is an area where consideration of wind condition may provide useful economic

benefits. If the regulators allow it, which may depend on flare quota considerations,

the actual production rate when the compressors were unavailable could be set based

on the measured windspeed and direction for the period in question. In other words

when the windspeed was low or in a beneficial direction the flaring rate could be set at

100% of production. Should this prove attractive to HMDC a set of envelopes for a

range of wind speeds and directions could be prepared which could be used in an

operational procedure to select production rate dependent on wind condition.


126

9.0 IMPLICATIONS FOR HIBERNIA

9.1 Introduction

In the foregoing sections the various aspects relating the RABS have been analysed.

The intent of this section is to combine the analysis into a form that can be used to

make decisions regarding potential capacity opportunities that exist in the flare system,

as well as identify the issues that will require resolution irrespective of the exercise of

any choices. Generally the potential changes fall into 3 categories:

Capacity opportunities resulting from the application of more modern design

practices (not all of these opportunities add apparent capacity).

Areas where the design documentation should be revised to increase the integrity

and traceability in the design.

Optional changes which can be considered to be related to house keeping.

Therefore this section is separated into three main sections; Firstly an outline of the

capacity opportunities is given including the apparent capacity effects the changes

would have; Secondly, a list and description of the important changes required to

ensure the integrity and traceability of the system design documentation is given;

Lastly, optional changes are described which will aid the future maintenance and

understanding of the design in future years.

Finally, a list of items that do not easily fall into the previous sections is included for

completeness.

In Appendix 2 a proposed scope of work is included which defines, in more

prescriptive terms, the work required to revalidate the flare system design assuming

HMDC decide to implement the changes described in this section in their entirety.

9.2 Flare System Capacity Opportunities

The following summarises the capacity opportunities available in the flare system with

respect to new codes and best practices.


127

Table 9.26 Effect of Changing Flare System Design Philosophy on the Apparent Design Capacity (Total Blowdown)

Issue Description /

concern

Case Meets current

code req's?

Safety Cost Failure potential Flare system

capacity*

Recommendation

Jet Fire Described in Section

6.2. Are vessels

sufficiently protected

from the effects of jet

fire?

Original design did not

purposefully include

mitigating measures

Design

Safety analysis not carried

through to engineering.

Codes do not

require

measures to be

included for jet

fire (API

currently

working to

change this)

N/A Rapid escalation. None Adopt best practice.

Ensure insulation

integrity on lower

pressure systems

during jet flame impulse

momentum.

Best practice

Detailed 3D analysis and

prevention measures to

ensure vessel will not fail in

a jet fire

Rapid escalation prevented

unless insulation fails.

Reducing

blowdown

start

pressure

Described in Section

6.3. Compressors

blowdown from PSHH

setting. Rest of

system depressures

from normal pressure.

Design Exceeds code

requirement

+ N/A N/A No effect on HP flare.

LP flare capacity

available is increased

by ~ 17,000 kg/h

compared to the

original blowdown case

89,601 kg/h

HMDC have declined

this change for now,

preferring the more

conservative design

approach (which avoids

changes to blowdown

calculations should

compressor operating

conditions change

significantly). The

capacity opportunity will

be described in an

appendix in the updated

RABS.

Best practice

As blowdown initiates

automatically, design

system with normal start

pressure

In the unlikely event that there

was a fire coincident with a shut

in situation (that was not caused

by trip) the blowdown rate would

be higher than anticipated and if

the wind were adverse could

lead to higher than planned

radiation levels on the platform.

Reducing

the


6.3 Certain vessels

Design Exceeds RP 521

requirements in

N/A N/A No significant effect on

LP flare because the

Best practice would

require all blowdown

8266-HIB-TN-C-0001 Page 128 of 139 Revision: B/tt/file_convert/55276c7749795994178b46d5/document.doc October 2000

Issue Description /

concern

Case Meets current

code req's?


capacity*

Recommendation

blowdown

end

pressure

(i.e. with wall

thicknesses over 1”)

can be depressured to

50% of the design

pressure rather than

690 kPag.

Various components are

depressured to either

690 kPag or 50% of the

design pressure

some instances. HP compressor seal oil

system requires the

compressor and

components to

depressure to

atmospheric pressure.

HP flare capacity

available is increased

by ~ 70% compared to

the original blowdown

case 133,316 kg/h

(reduces to

~40,000 kg/h)

calculations to be re-run

and new orifice plates in

the affected blowdown

section. For the

moment best practice is

declined. An appendix

to the updated RABS

will be created to

identify the capacity

opportunity in case it is

required in the future.

Best practice

Systems with design pressure above 1724 kPag should be depressured to 50% of

the design pressure. Systems with design

pressure below

1724 kPag need not be

depressured. However, if

it is chosen to do so, the

final pressure should be

690 kPag or 50% of the

design pressure,

whichever is less.

Vessels with wall

thicknesses below 1 inch

should be considered

separately.

More closely

follows the intent

of RP 521

+ N/A

Blowdown

stagger


6.4. Stagger not

sufficiently

independent and

equipment is in same

fire zone. Fire may

affect A train injection

compressor, yet

blowdown will be held.

Design Ambiguous

(although the

recommended

practices allow

controlled

blowdown)

1. Stagger fails closed would

lead to escalation.

2. Stagger fails open at

initiation leads to high radiation

levels.

3. Jet fire analysis suggests

injection compressor vessels

should not fail.

33,974 kg/h of load is

added to the LP flare

blowdown case.

(compared to

89,601 kg/h original LP

flare blowdown design

case. New rate is

therefore

Decline best practice.

The A train injection

compressor

components are not at

risk due to their

thickness.

QRA the software

reliability.


Issue Description /

concern

Case Meets current

code req's?


capacity*

Recommendation

123,575 kg/h).Best practice avoids stagger

unless the systems can be

made sufficiently

independent

+ 1. Difficult problem relating to

back pressure on LP separator

to overcome.

2. Otherwise, very reliable.

Higher radiation levels designed

for.

Design

windspeed

and direction


6.6. The design

windspeed is higher

than absolutely

necessary.

Design = 27 m/s from North No code

requirements.

N/A If windspeed is higher and

design release is occurring the

radiation levels on the platform

will be exceeded.

HP and LP flare

apparent capacity is

increased by

approximately 7% (i.e.

by 9,000 kg/h and

6,000 kg/h

respectively).

Best practice declined.

For continuous flaring

case a risk mitigation

procedure could be

developed to increase

the flaring rate when the

compressors were down

dependent on the

measured wind

condition.

Best practice = 24.5 m/s

from North

34.2 m/s from North West

No code

requirements

If windspeed is higher and

design release is occurring, the

radiation levels for emergency

on the platform will be exceeded

somewhat. It is highly unlikely

that anyone would be on deck

without the necessary protection

in such a case.

Acceptable

flare

radiation

levels


6.7. The requirements

in the RABS are over

conservative.

Design

6.3 kW/m2 at crown block and 3.2 kW/m2 at escape ways

Over

conservative

N/A N/A HP and LP flare capacity is increased by approximately 50% before new radiation levels are approached during blowdown (i.e. by 60,000 kg/h and 45,000 kg/h for the HP and LP flares respectively).

Adopt and describe best practice in flare documentation. The practice will remove inconsistency compared to Canadian regulations and international standards. Hydraulic considerations may not allow the full use of new capacity.

Best practice - Radiation levels raised to:

9.5 kW/m2 at crown block (shielded)

6.3 kW/m2 at any escape way (no shielding)

3.2 kW/m2 at the helideck

1.9 kW/m2 continuous at the weather deck

(and meets

Canadian

regulations)

N/A


Issue Description /

concern

Case Meets current

code req's?


capacity*

Recommendation

Incorporate

the effects of

vessel

insulation on

the vapour

rates


6.9.1.

Design

No credit taken for insulation on vessels in blowdown calculations

Over

conservative

N/A N/A HP and LP flare capacity is increased by approximately 5% during blowdown case (i.e. by 6,000 kg/h and 4,000 kg/h for the HP and LP flares respectively).

Best practice would require all blowdown calculations to be re-run. For the moment best practice is declined. A note will be incorporated in the revised flare documents to note the capacity opportunity in case it is required in the future.

Best practice

Credit taken for insulation

N/A

Remove the

effect of

future

equipment


7.0.

N/A N/A N/A N/A N/A HP flare capacity is increased by approximately 38,000 kg/h during blowdown. LP flare system capacity unchanged.

Incorporate the revised data in the updated RABS. Identify the “spare” capacity for future projects in a suitable section.

N/A N/A N/A Negl. N/A

= acceptable - + = most acceptable

= least expensive - = most expensive

* by making change to best practice


Care needs to be exercised when considering Table 9.26 as the values are not

additive. What the table does show however is the significant spare capacity in the HP

flare when considering the blowdown cases. The LP flare is very different. The

changes that affect capacity alone (rather than implied through radiation calculations)

are insufficient to offset the large change required to remove the blowdown stagger.

Therefore this change would force the design rate of the system to be increased and

would therefore require detailed hydraulic analysis to be undertaken. The difficult

aspect will be the superimposed back pressure on the LP separator. If this is too high

there will be the undesirable consequence of raising the pressure in the LP separator

when the blowdown valve opens. This would have the effect of increasing inventory

and reducing the time to failure of the LP separator if exposed to fire. Some mitigating

measures would likely be required in this instance. Given this and the apparent

inherent capability of the injection compressor components to survive the pause period

before blowdown commences, suggests that the stagger in the system should be

retained.

For the relief cases other than blowdown, two out of three of the capacity opportunities

(i.e. relief cases which were overestimated during design) are negative. The worst of

the problems relates to the spillover valve failure cases as these have the potential to

significantly exceed the HP flare system design rate. This is shown on Table 9.27.

which follows.


Table 9.27 Effect of Changing Flare System Design Philosophy on the Apparent Design Capacity (Relief Cases)

Issue Description /

concern

Case Meets current

code req’s?

Safety Cos

t

Failure potential Flare system capacity* Recommendation

Two-phase

relief (See

Section 5.3

and 6.5)

New sizing

method

increases valve

size required for

this case

Design - Old additive API method No longer N/A If the design case is the

sizing case the vessel can

be overpressured.

The design rate which can

be accommodated in the

existing valves reduces

dramatically. For

comparison max single well

rate for HP separator is ~

54 kbopd.

Adopt best practice.

This issue requires the

maximum well rate and

maximum number of

wells to be redefined as

they are likely to

compromise the RV size

on the HP separator.

The test separator RVs

are being replaced.

Best practice - New API method

Valve size may be to high

and valve will chatter.

Missed relief

cases (See

Section 5.2)

Failure of

spillover valves

(open) exceeds

flare system

capacity.

Design - missed a valid relief case No. N/A If valve fails open the flare

system design rate is

significantly exceeded.

Flare system capacity was

not dimensioned for the

dimensioning case.

Adopt best practice and

use measures to limit

peak load.

Best practice - Design for any single

valve failure.

N/A

Blowby

cases

methodology

flawed (See

Section

5.3.1)

Blowby cases

are over

conservative.

This would

prevent the

installation of

larger LCVs if

this proved

necessary.

Design - Over conservative case

assumed.

N/A Valve is likely to chatter if

faced with the blowby

case.

As there will be no desire to

change the existing valve,

there will be a latent capacity

in the system which can be

used for future upgrades.

The capacity change

available (which would

translate to an increase in

separator LCV size) is

approximately 20%.

Add note to RABS

update to describe the

spare capacity.

Best practice:

Use settleout pressure for

shutdown case

Take account of downstream

control valve positions and fluid

properties in production case.

N/A A valve, properly sized for

the case in question, will

not chatter when faced

with the design case.

= acceptable - + = most acceptable

= least expensive - = most expensive

* by making change to best practice


In the above tables the issues which add capacity are optional. In other words these

are issues that HMDC can adopt or decline at will. The issues which have a negative

capacity effect, for obvious reasons, will require some work to resolve.

9.3 Impact on the Design Documentation

The outcome of the technical audit indicates the following aspects of the design will

need to be revisited / revised. The table is a significantly shortened version of the

table presented in Section 5.4 and represents the most important changes that should

be considered. Repetitive items (including those in tables 9.1 and 9.2) and issues

requiring simple comment in the RABS update are omitted. The table should be read

in this light.



Number

34-

Title Number

34-

Description Action




calculations


used?







Update RABS.





See also 34.010/1.


cooldown before

hydrate formation &

minimum

temperatures



sections












KO Drum size







platform?








Separator and MP to

LP Separator Blowby

Cases






design.






- Max Relief Case)





capacity.

025 / C 3rd Stage

Compressor Max


Analysis










Separator Heaters

Simultaneous Fire

Relief - Network

Analysis




Construct an LP flare network







31.37 Relief Valve

Calculations - LP

Separator




pressure mode?




this.

In Granherne’s opinion none of the above items are optional.

9.4 Optional Changes

The following changes could be considered to increase the integrity and traceability of

the design work.

Table 9.30 Technical Audit Optional Changes (System 34)

Number

34-

Title Number

34-

Description Action

045 / E Total HP Blowdown

Initial Conditions

(Checks blowdown

line sizes for

individual system

blowdowns)

045/1 There is no network analysis run with

common HP Blowdown at initial

conditions

Consider constructing a HP flare

network model to assess future

modification projects against.

045/2 Consistency error in the number and

flows in the gas injection flowlines

Add a note to the RABS clarifying

the injection manifold rate basis.

046 / G Fuel Gas Cooler /

Heater tube rupture


046/1 ''As Built' P&IDs show bursting discs in

this service (calc considers PSVs)

therefore calc is no longer valid

There is no replacement

calculation for the installed

bursting discs. The bursting disk

calculations should be reviewed to

identify implications for the flare

system.

050 / G 3rd Stage Suction

Scrubber A (D-3303A)

PSV Discharge Line

Size Confirmation


pressure = 8200 kPa(g), 'As Built' P&ID

shows set pressure = 7000 kPa(g)

P&ID set pressure error?

059 / G Comparative Program

check of INPLANT

Single Phase

Simulation vs ESI

059/1 Accuracy of calculations using ESI

instead of INPLANT

Revisit ESI calculations and

replace as necessary


Table 9.31 Technical Audit Optional Changes (System 31)



31.36/6 Are the gas blowby cases are

methodologically flawed?

Add note to RABS update

31.38 Inlet Line Size

Checking for Relief

Valves

31.38/1 Inlet line sizes should have been

recalculated using 'Final' relief data

and isometrics.

Check / redo inlet line sizing

calculations as necessary.

These items are optional as the inconsistencies are minor.

The other area that could be improved is the overall level of as built of the design

calculations. Generally the calculations were never revised for key design data late in

the project. This included the calculation of inventory and the use of vendor supplied

settleout pressures. The latter item becomes more important if the changes above are

pursued. When the relief valve and control valve data sheets were prepared

superseded design data was also used. Our initial analysis of these combined effects

suggests that they are benign. For example we expect, but cannot be sure, that the

inventory assumptions will actually be conservative; Our analysis of the valve

calculations shows that in some cases the valves may have ended up being slightly

smaller than desired but if this were a problem it would have shown up by now.

This type of inconsistency is relatively common; there is never the perfect design

project. HMDC will need to decide whether they can accept the inconsistencies.

9.5 Miscellaneous Requirements

9.5.1 Updating the Design Documentation

The above changes generally point to a requirement to rerun some of the key flare

system design calculations. Whether or not this uncovers areas where hardware

changes will need to be made will have to be seen; nevertheless there will still be the

requirement to update the design volumes and make the calculation changes

traceable. This will be in addition to making the new calculated information obvious in

the updated RABS.


The following summarises the type of changes to the design documentation that will

be necessary:

Relief and Blowdown Study Report

Fairly extensive rewrite of the report.

Design Calculations

For each change prepare a calculation revision which revs up the existing

calculation (in other words building on the existing work). This would include:

- Calculations identified above.

- Flare radiation calculations (for windspeed and allowable radiation levels)

- Continuous radiation cases. Analysis of allowable production rate vs wind

speed.

Blowdown inventory calculations (for removed inventory)

Reliability analysis of the system that controls the compressor stagger, to ensure

the system is sufficiently reliable to ensure the design integrity.

9.5.2 Implementation Projects

In this section there are some projects mentioned which will in all likelihood require

hardware changes to be made (resulting from the above there may be more).

Insulation conformance - The explicit ability of the platform to cope with a jet fire

hazard requires the insulation around the vessels to remain in place during jet

flame impingement. This may require the insulation strength to be improved.

Modifications to limit peak flaring rate during spillover valve failure.

Instrument modifications to warn operators when the requirement to blowdown

compressors is becoming imminent (to avoid low temperature problems).


10.0 REFERENCES

1. Canada-Newfoundland Atlantic Accord Implementation Act - Newfoundland

Offshore Petroleum Installations Regulations, February 21,1995.

2. Concept Safety Evaluation, Cremer and Warner, October 1991.

3. Fire Risk Analysis, Cremer and Warner, May 1992.

4. Fire Risk Analysis Update, Cremer and Warner, February 1993.

5. Design Phase Risk Assessment, Caldwell Consulting, May 1995.

6. Design Phase Safety and Environmental Assessment, Doc No. CM-E-F-R-M00-

RP-104 Rev C0, May 1995.

7. Structural Passive Fire Protection Analysis, Aker Engineering, 21 February 1993.

8. Review of Emergency Systems for the Proposed Hibernia Platform, Cremer and

Warner, Report No. 93432, 15th March 1994. HMDC Doc No: CM-Y-F-R-M00-RP-

008.000 Rev 001.

9. Deleted.

10. Deleted.

11. Gayton P.W. and Murphy, S.N. (1995) Depressurisation System Design, IChemE

workshop, “The Safe Disposal of Unwanted Hydrocarbons”, Aberdeen 1995.

12. Deleted.

13. Kaldair Design Data Dossier, Doc. No. CM-Z-M-Z-210-ZM-1083-002.0, December

1995.


APPENDIX I

CALCULATION TECHNICAL AUDIT

8266-HIB-TN-C-0001 Appendix I Revision: B/tt/file_convert/55276c7749795994178b46d5/document.doc Page 1 of 42 October 2000


Calculation Description Rev DateNumber /

Calculation Book

34-005 / A Blowdown Section Inventory Calc 06 18-May-93

(Provides input to blowdown simulations)

Audit Tasks Methodology X see /2&4 Consistency As Built X see /1,3&5

Key Assumptions

Blowdown volumes calculated using piping isometrics available at that time. Not as built.

Included between 10 - 20% margin for small bore pipework.

Future Gas Dehydration Unit blowdown section included

Future MP Manifold blowdown section included

Future Gas Lift Manifold blowdown section included

Not all BD volumes recalculated using final piping isometrics

Vessel weights estimated?

Blowdown of GT driven compressors must be to atm conditions in 15 mins (seal oil system rundown time),

therefore based on estimated settle out pressure + 5% contingency on section vol. The vessels were assumed on fire.

1st Stage Compressor uses dry gas seals therefore not as critical as above

PIDs at Rev C2, Isometrics at Rev B

Key Results

Results Summarised in table below

Section Total Volume Liquid Charge Wetted Area Section Weight Initial Pressure1 Initial Temp1

m3 m3 m2Tonnes kPa(a)

oC

HP Separator 186.81 72.70 98.40 165.00 - -

MP Separator 340.60 78.07 121.80 175.00 - -

West Test Separator 48.53 11.30 30.60 45.00 - -

Future Test Separator 33.50 6.96 29.04 36.00 - -

East Test Separator 49.05 9.66 28.11 46.50 - -

HP Manifold 54.50 (2) 200.00 234.91 - -

MP Manifold 43.45 (2) 125.00 193.13 - -

Test Manifolds 2.70 (2) 50.00 11.69 - -

HP Offgas Manifold 16.97 - - 32.00 - -

MP Offgas Manifold 17.60 - - 23.87 - -

Gas Injection Manifold 4.00 - - 40.00 - -

Gas Injection Flowline 1.50 - - 40.00 - -

Future Gas Lift Manifold 1.70 - - 17.21 - -

Future Gas Dehydration Unit

8.38 - - 50.00 - -

2.55 - - 10.00 - -

1st Stage Gas Compressor 55.62 5.00 10.80 35.00 316 91.0

2nd Stage Gas Compressor 24.92 (2) 10.20 45.00 2,046 80.0

3rd Stage Gas Compressor 19.18 (2) 6.60 75.00 6,122 86.8

4th Stage Gas Compressor 9.15 - 100.00 24,115 117.0

Notes

1. Values only given for compressor sections where settle out conditions calculated

2. Not required as input as auto calculated by blowdown program

Issues

34-005/1 Are the blowdown volumes used sufficiently accurate?

34-005/2 Jet fire scenario not taken into account for the design of the blowdown system

34-005/3 Were the real settle out pressures ever used?

34-005/4 Were fire areas used for total blowdown rate?

34-005/5 Are vessel weights used reasonable?

Audited by AJR Date 16-Jun-00 Checked by MFG Date 15-Aug-00

Gas Dehydrator, Inlet Cooler & Scrubber

Gas Dehydrator Discharge Cooler & Scrubber



Calculation Book

34-005 / A Blowdown Section Inventory Calc 05 08-May-92

(Provides input to blowdown simulations)

Audit Tasks Methodology Consistency As Built X see Rev 06

Key Assumptions

See 34-005 Rev 6 and 34-006 for final data used in blowdown analysis.

Key Results

See 34-005 Rev 6 and 34-006 for final volumes used.

Issues

See 34-005 Rev 6 and 34-006.




Calculation Book

34-006 / A Blowdown Summary 05 19-May-93

Audit Tasks Methodology X See /2-4 Consistency X See /1 As Built X See /5

Key Assumptions

Blowdown model isentropic efficiency set at 0.5

Blowdown from PSHH to LP flare and from normal operating pressure to HP flare.

Assumes heat input from fire.

Constant rate blowdown from LP Separator and FG Separator (500 and 250 kg/h respectively).

Uses estimated settleout pressures?

Key Results

Rev 04 calculated peak values remain the basis for flare spec (Rev 05 calc provides final summary of BD loads)

Rev 05 calculated values not incorporated into Relief & Blowdown Study Report Rev C2

Total LP Blowdown Flows

- Minor changes for peak total LP BD flows between rev 04 & 05 (though individual flowrates have changed)

- Basis for spec of total LP BD load to vendor considers blowdown of GT driven compressors from PSHH settle out conditions

in all stages for both trains. This case is considered of low probability

- Sensitivity check on blowdown from compressor settleout pressure. Depressuring from normal operating pressure

increases apparent system capacity by 19%

- Total peak BD flow (staggered from PSHH settle out conditions) for this calc rev 86,709 kg/h [89,601 kg/h in Relief & Blowdown

Study Report] - Detailed in table below

Total HP Blowdown Flows

- Total peak HP BD load increase by 5.8 % since rev 04 calc

- Includes for future / possible future equipment including: MP Manifolds / Future Test Separator & Manifolds / System

Gas Lift Manifolds / Gas Dehydration

- Future Gas Dehydration unit is significant (20%) proportion of total HP BD load. The volume used in BD calc is approx double

expected value

- Total peak BD flow this calc rev 141,335 kg/h [133,616 kg/h in Relief & Blowdown Study Report] - Detailed in table below

Design Case Blowdown Summary Table - Initial Blowdown Rates (time = 0)

System Initial Blowdown Rate, kg/h System Initial Blowdown Rate, kg/h

Inj Comp 'A' 0 Test Manifolds 588

Inj Comp 'B' 36,690 HP Manifold 3,239

3rd Stage Comp 'A' 15,917 MP Manifold (future) 999

3rd Stage Comp 'B' 15,917 Test Separator E & W 23,559

2nd Stage Comp 'A' 7,325 Test Separator (future) 8,262

2nd Stage Comp 'B' 7,325 HP Separator 41,311

1st Stage Comp 2,785 MP Separator 8,284

LP Separator 500 GI Manifold 4,154

LP FG KO Drum 250 GL Manifold (future) 3,075

FG Cooler 681

HP FG KO Drum 6,093

Offgas Manifolds 4,096

Dehydration System (future) 28,062

GI Flowlines 8,932

Total 86,709 Total 141,335

1. Blowdown from PSHH Settleout conditions, 3 minute time delay on Inj Comp 'A' blowdown

Issues

34-006/1 HP Blowdown calculation higher than vendor aware of. Radiation level for case is underestimated.

34-006/2 Correct isentropic efficiency used?

34-006/3 Is design case too extreme?

34-006/4 Is constant rate blowdown a valid design method, i.e. not according to API?

34-006/5 'As Built' settleout pressure


Blowdown to HP FlareBlowdown to LP Flare1



Calculation Book

34-006 / A Blowdown Summary for RFD Flare Report 04 14-Feb-92

Audit Tasks Methodology X see Rev 05 Consistency As Built X see Rev 05

Key Assumptions

Superseded by Rev 5

Key Results

Total LP Blowdown Flows

- Total peak LP Flare BD flow for this calc rev as per Relief & Blowdown Study Report Rev C1, i.e. 89,601 kg/h (taking into account

3 minute stagger on Inj Compressor 'A' blowdown) - Detailed in table below

Total HP Blowdown Flows

- Total peak HP Flare BD flow for this calc rev as per Relief & Blowdown Study Report, i.e. 133,611 kg/h - Detailed in table below

Design Case Blowdown Summary Table - Initial Blowdown Rates (time = 0)

System Initial Blowdown Rate, kg/h System Initial Blowdown Rate, kg/h

Inj Comp 'A' Test Manifolds

Inj Comp 'B' HP Manifold

3rd Stage Comp 'A' MP Manifold (future)

3rd Stage Comp 'B' Test Separator E & W

2nd Stage Comp 'A' Test Separator (future)

2nd Stage Comp 'B' HP Separator

1st Stage Comp MP Separator

LP Separator GI Manifold

LP FG KO Drum GL Manifold (future)

FG Cooler

HP FG KO Drum

Offgas Manifolds

Glycol Column (future)

Glycol Discharge Scrubber (fut)

Total 89,601 Total 133,616

1. Blowdown from PSHH Settleout conditions, 3 minute time delay on Inj Comp 'A' blowdown

Issues

See 34-006 Rev 5



Calculation Book

34-008 / A Flare gas LHV calc for RFD flare report 03 13-Apr-92

Audit Tasks Methodology Consistency As Built Key Assumptions

N/A. Based on design compositions.

Key Results

LHV in Flare Data Sheet

Issues

None


Blowdown to LP Flare1 Blowdown to HP Flare

8,484

8,484

0

33,974

250

4,188

5,250

3,589

20,098

4,047

50,244

8,840

17,562

17,562

1,117

21,969

6,093

2,832

3,075

617

1,657

2,785

500



Calculation Book

34-009 / A Metal Surface Temperature Calcs for AFD Flare Report 03 20-Apr-92


Metal temperature model derived from Kern's Process Heat Transfer and Kent's Mechanical Engineers Handbook.

Ambient Air Temp = 20C

Emissivity = 0.7

Windspeed = 0 / 27 m/s

Pipe Flame Buoyancy = 3.0 m/s

Sonic Flame Buoyancy = 4.6 m/s

Stack Length = 115m

HP Flare Tip angle 45 deg to Horizontal

LP Flare Tip angle 90 deg to Horizontal

Windspeed = 0 / 27 m/s

Calorific Value = 47 MJ/kg

Key Results

Metal Temperatures at the following point locations considered (m):

North East Elevation Worst Case Temp, oC

Top Weather Deck 75 0 30 70.3

Crown Block 55 10 93 109.0

Finger Board 58 6 69 93.2

Platform Crane 84 42 49 87.0

Boom Base 75 0 0 55.8

30m up Flare Boom 90 0 26 72.7

50m up Flare Boom 100 0 43 91.8

70m up Flare Boom 110 0 61 127.2

90m up Flare Boom 120 0 78 198.0

105m up Flare Boom 128 0 91 335.4

110m up Flare Boom 130 0 95 420.5

112.5m up Flare Boom 131 0 97 482.7

Issues

None.

Temperatures are for information and paint selection. Not dimensioning design criteria.



Calculation Book

34-009 / A Metal Surface Temperature Calcs for AFD Flare Report 02 16-Jan-92


As Rev 03 of this calculation

Key Results

Superseded by Rev 03 of this calculation

Issues

As Rev 03 of this calculation




Calculation Book

34-010 / A Calculation of allowed cooldown before hydrate formation & minimum 02 23-Mar-92

temperatures achieved in flare gas from critical blowdown sections

Audit Tasks Methodology Consistency As Built X See /1

Key Assumptions

Ref. Process simulations: HIBBD126A.LIS, 126B, 127, 137, 138, 139, 159, 160, 162

Key Results

1. Injection Compressor Blowdown Section

Initial settle out conditions = 22,669 kPa(a), Temp = 108.2 C

Hydrates form if gas cools to 61 C before blowdown commences

2. 3rd Stage Compressor Blowdown Section



3. 2nd Stage Compressor Blowdown Section



4. Inlet Gas Manifold Blowdown Section

Initial conditions = 40,000 kPa(a), Temp = 156 C

Hydrates form if gas cools to 67.5 C before blowdown commences

Issues

34-010/1 Was the calculation methodology sufficiently robust?

34-010/2 Should 'troubleshooting' methanol injection points be incorporated?



Calculation Book

34-011 / A Review of HP flare KO Drum size 03 09-Mar-92


Max HP relief case = 244,897 kg/h, MW = 20.65, dP(tip) = 500 kPa, T = 59.0 C, z = 0.98, k=1.24 (as per R & BD Study Report Rev C1)

Calc'd KO Drum operating pressure (739 kPa(a)) based on estimated pipe equivalent length (drum to tip)

Droplet size 400m

Drum sizing to API 521

Vol liquid required to hold = 37.18 m3 (for basis see 34-011 Rev 02)

Key Results

HP Flare KO Drum reduced to 2.8m Dia x 7.5m T/T

Issues

None




Calculation Book

34-011 / A Review of HP flare KO Drum size 02 06-Feb-92

Audit Tasks Methodology Consistency X See /1 As Built X (See 34-011 Rev 03)

Key Assumptions

Liquid volume sizing basis: 10 mins relief of 1 well at max flow and 1 well at average flow (i.e. 10 mins at 40,000 bpd)

Overall sizing basis: Drum at maximum liquid level + Max HP flare relief load (274,878 kg/h, MW = 21.8, T = 83.7 C

z = 0.98, P = 743 kPa(a))

Key Results

HP Flare KO Drum reduced to 2.8m Dia x 8.0m T/T (Superseded by Calc 34-011 Rev 03)

Issues

34-011/1 A note on the front of calc 34-064 states that Rev 7 of Design Basis gives max well flow of 20,000 bpd + average

well of 10,000 bpd, i.e. 30,000 bpd total. The individual well design rate has changed. What are the implications

for the platform?



Calculation Book

34-012 / A Review of LP flare KO Drum size 03 10-Mar-92


Max LP relief case = 110,874 kg/h, MW = 25.37, dP(tip) = 10 kPa, T = 67.5 C, z = 1.0, k=1.19 (as per R & BD Study Report Rev C1)

Calc'd KO Drum operating pressure (147.7 kPa(a)) based on estimated pipe equivalent length (drum to tip)

Droplet size 400m

Drum sizing to API 521

Liquid transfer from HP Flare KO Drum = 37.13 m3

Key Results

LP Flare KO Drum reduced to 2.8m Dia x 8.0m T/T

Issues

See 34-011/1



Calculation Book

34-013 / A Preliminary calc of HP Flare tip delta P vs flowrate 01 11-Jun-91


Superseded by vendor (Kaldair) supplied curve

Vendor data based on HP flowrate of 244,897 kg/h, MW=20.65, T=59 C

Key Results

Superseded by vendor (Kaldair) supplied curve

Issues

None




Calculation Book

34-014 / A HP Flare Drum Pump Calculations 02 02-Jul-91


HP Flare Drum size is 2.8m Dia x 8.5m T/T

Total drum volume is 58.05m3

Pump capacity designed to Mobil E&P661 (Empty half drum volume in 2 hours)

Key Results

15m3/h pump capacity

Issues

None however 'As-Built' HP Flare Drum size reduced to 2.8m Dia x 8.0m T/T



Calculation Book

34-015 / A Calc to review options for reducing HP to MP Separator and 01 14-Aug-91

MP to LP Separator Blowby Cases

Audit Tasks Methodology X See /2 Consistency X See /1 As Built (See Note)

Key Assumptions

Maximum allowable LP Flare flowrate = 119,324 kg/h (Higher than design rate stated in R & BD Study Report Rev C1)

Maximum allowable HP Flare flowrate = 274,878 kg/h (Higher than design rate stated in R & BD Study Report Rev C1)

Uses Masoneillan sub-critical flow correlation except for HP to MP case with 2 off control valves (critical flow)

Key Results

For MP to LP Separator blowby relief rate not to exceed maximum allowable LP Flare load:

Calculates min LP Sep design pressure of 9.2 barg with single valve Cv=1300 , or

Original LP Sep design pressure (7 barg) but 2 off control valves (with independent transmitters and controllers) Cv=700 each

For HP to MP Separator blowby relief rate not to exceed maximum allowable HP Flare load:

Calculates min MP Sep design pressure of 33.5 barg with single valve Cv=750, or

Original MP Sep design pressure (20 barg) but 2 off control valves (with independent transmitters and controllers) Cv=375 each

Issues

34-015/1 Relief & Blowdown Study Report Rev C1 non-concurrent maximum allowable LP and HP Flare loads are

110,874 kg/h and 244, 897 kg/h respectively. Rates used in these calculations exceed design.

34-015/2 Is considering only one control valve fails open for gas blowby case when 2 installed in parallel

realistic / allowable even with provision of independent transmitters and controllers?

Note: Actual installed control valve CVs (PID for HP Separator, Rev E2, states max control valve CV=350, datasheet states CV=330)

(PID for MP Separator, Rev E3, states max control valve CV=650, datasheet states CV=600)




Calculation Book

34-016 / A Flare purge flowrate calc 03 18-Feb-93


Sweep velocity 0.2 m/s

Sweep gas MW = 20.21

Sweep gas Temp = 10 C or 37 C

Lines Swept HP LP

16"-VH-34153 14"-VL-35299

12"-VH-34278 20"-VL-35263

14"-VH-34208 14"-VL-35261

16"-VH-34156 14"-VL-35241

14"-VH-34157

14"-VH-34151

Key Results

Purge gas mass flowrates for flowmeter specification

Issues

None



Calculation Book

34-017 / A Calculation of flare gas Peak Velocity at the Flare Gas Flowmeter Location 01 21-Nov-91

Audit Tasks Methodology Consistency X (Out of date max flows)

As Built

Key Assumptions

Max HP Flare Gas Rate = 274,878 kg/h (giving vel = 91.73 m/s) [18" N.B. Line]

Max LP Flare Gas Rate = 119,324 kg/h (giving vel = 89.23 m/s) [24" N.B. Line]

Max LP Flare Gas Rate (new design value) = 131,473 kg/h (giving vel = 92.23 m/s) [24" N.B. Line]

Key Results

Maximum gas velocities as given above

Issues

None as calc was only developed to assist ESIN with flowmeter evaluation




Calculation Book

34-018 / A Comparison of Flare Vendor Radiation Levels 02 30-Jan-92


Windspeed = 0 or 27 m/s

Flame Emissivity Values:

Birwelco calculations include for solar radiation

Boom Length assumed 123 m

Flare Tip angle 45 deg

Key Results

Radiation levels at key points on the platform for the different flaring cases.

Issues

None.

This calculation was used for selection only. Calculation was eventually superseded by the flare vendor's 'As Built' data package.

See also calculation 34-062



Calculation Book

34-019 / B Line sizing calculations for lines discharging to HP & LP Flare systems 01 30-Jan-92

(i.e. Relief, Blowdown & Spill-off valves)

Audit Tasks Methodology Consistency N / A As Built N / A

Key Assumptions

Preliminary line sizing calculation for input into deck level flare headers and lines upstream of blowdown valves only

Lines where velocity is greater than 0.35 Mach identified as requiring angled entry into flare header

Key Results

Calc is superseded by Network Analysis runs.

Issues

None


Case HP Flare LP Flare

Kaldair Birw elco Kaldair Birw elco

1 0.1 0.121 0 .18 0.148

2 0.1 0.107 - -

3 0.1 0.122 - -

4 - - 0.2 0.3

5 0.1 0.124 0.2 0.3



Calculation Book

34-020 / B LP Flare KO Drum Level Valve Calculation 01 30-Jan-92

Audit Tasks Methodology Consistency As Built (See Issues)

Key Assumptions

Masoneillan sub-critical flow formulae

Installed CV (LV-0009) = 65 (P&ID shows valve number to be LV-0109)

4" liquid outlet line

Key Results

Line size and installed control valve CV are adequate

As a result of calc LP Flare KO Drum heater duty revised to 30kW

LSLL set point revised to ensure drum heater element is always covered

Issues

None for this calc (See 34-021 below). Control valve datasheet confirms actual 'As-Built' CV=70



Calculation Book

34-021 / B Gas blowby calc from LP Flare KO Drum to open Hazardous Drains Tank (F-5201) 01 06-Feb-92

& Calc of F-5201 Req'd Vent Line Size (Result of HAZOP comment)


Key Assumptions

LP Flare KO Drum at its maximum pressure (i.e.135 kPa(a) )

Masoneillan sub-critical flow formulae

Installed CV (LV-0009) = 65 (P&ID shows valve number to be LV-0109)

Key Results

Built-up back pressure at F-5201 approximately 1.1 kPa(g). (Design pressure on 'As-built' P&ID = Static head + 3 kPa(g).

Static head calculated to be 20 kPa based on tank water full to overflow level)

Issues

None. Actual 'As-Built' control valve CV=70 however calculation results will not change significantly.

Vessel number is 'F-5210 on 'As-Built' P&IDs.




Calculation Book

34-060 / B Indicative Injection Compressor Cooldown Calculation 01 29-May-92


Key Assumptions

Internal B&R program 'ColdSpot' used, system (incl compressor, scrubber etc) converted to single pseudo pipesize for calc purposes

Injection Compressor settle out conditions P=22669 kPa(a), T=108.2C

HT coefficients based on either natural convection + ambient temp of -10C or forced convection (wind speed 30ft/s)

Min flare temp = -45C

Key Results

Cooldown temp at which hydrate formation occurs in LP flare system on section blowdown = 61.5 C

Hold time for cooldown temperature to reach 61.5 C, hours: 3.4 (No insulation, natural convection)




Cooldown temp at which minimum design temperature occurs in LP flare system on section blowdown = 50 C

Hold time for cooldown temperature to reach 50 C, hours: 4.5 (No insulation, natural convection)




Results support philosophy that compressor sections will not remain isolated at pressure for periods in excess of 1-2 hours

Issues

See 34-010/1 and 34-010/2



Calculation Book

34-061 / B Simplistic Steady State Preliminary Review of the Annulus Rupture Relief Flowrate 01 10-Sep-92


Key Assumptions

Based on 4" choke valve with max CV = 174

Two possible sources of Lift Gas

- 3rd Stage Compressor (P=18,237 kPa(a), T=177.4 C, MW=21.86, z=0.937)

- Future Lift Gas Dehydrator (P=13,700 kPa(a), T=38 C, MW=22.09, z=0.66)

Lift gas supplied at 13,000 kpa(a) downstream of choke valve

Relief flow based on Masoneillan critical flow formulae

Critical flow factor = 0.82

Key Results

Relief flowrate based on 3rd Stage Compressor 228.6 MMSCFD

Relief flowrate based on Future Lift Gas Dehydrator 319.6 MMSCFD

Max HP flare design load 237.85 MMSCFD

Installed choke valve CV must be around 120 if lift gas supplied from Future Lift Gas Dehydrator to avoid exceeding flare design load

Issues

34-061/1 Annulus rupture case had the potential to be the defining case for the HP flare system (depending on installed

choke valve CV). What happened subsequently?




Calculation Book

34-062 / B Sensitivity review of max emergency radiation levels on total blowdown case 01 21-May-93

if injection compressor 'A' staggered blowdown mechanism fails


Calc for indicative only, design values by flare vendor

Wind speed = 27 m/s

3 cases considered, total platform blowdown with:

1 - Turbine driven compressors at PSHH settle out conditions, staggered blowdown (3 mins)

2 - Turbine driven compressors at normal settle out conditions, staggered blowdown (3 mins)

3 - Turbine driven compressors at normal settle out conditions, staggering mechanism fails

HP tip at 45o angle

Stack length = 115 m

Key Results

Case 1 Case 2 Case 3

Location

Crown Block 7112 (2254) 6029 (1911) 8104 (2569)

Finger Board 5094 (1615) 4490 (1423) 5644 (1789)

Weather Deck 3013 (955) 2718 (861) 3284 (1041)

Reconfirms requirement for staggered blowdown to meet radiation specs however increased radiation

deemed acceptable for short periods

Issues

None. Case 2 results were superseded by vendor calculations. The vendor calculations confirm the staggering requirement.

The calculations used earlier tip orientations, however as the Case1 and Case 3 calculations are indicative only this is

not considered significant.



Calculation Book

34-063 / B Summary of Tube Rupture Relief Flare Line Liquid Velocity 01 01-Jun-93

(Calcs to Provide Information Requested by Piping Stress)


No details of method of tube rupture flowrate given in this calc

Key Results

Fluid velocities (for stress calculation purposes)

Issues

None


Radiation Levels, W/m2 (Btu/ft2/h)



Calculation Book

34-022 / C HP Flare Network Sizing 02 22-Mar-93

(HP Separator - Max Relief Case)

Audit Tasks Methodology Consistency X See /1 As Built X See /2

Key Assumptions

HP Flare system iso rev A4

HP Separator Blocked Outlet (227,649 kg/h total load)

PSV Datasheet at Rev C1 indicates multiple PSVs with staggered set pressures on HP Separator relief valves. 3 x 50% PSVs,

(2 operating + 1 standby). 2 are set at vessel DP + 1 set at 105% DP

Blocked outlet relief is 100% vapour (no liquid or 2 phase relief)

Tip P estimated from Kaldair supplied graph

Maximum pressure at PSV discharge (Normal + Built-up back pressure) = 851 kPa(a)

Key Results

Maximum pressure at PSV discharge for this case = 943 kPa(a)

Maximum Mach No. in system = 0.25

Calculated tip P = 415 kPa

Issues

34-022/1 Calculated maximum pressure at PSV discharge exceeds value on PSV datasheet Rev C1

34-022/2 Effect of increased production / production fluid GOR



Calculation Book

34-023 / C HP Separator Max Spill-off Case - Network Analysis 02 22-Mar-93

Audit Tasks Methodology X See /2 Consistency X See /1 As Built Key Assumptions


HP Separator Max Spill-off (244,897 kg/h total load)

Max spill-off relief is 100% vapour (no liquid or 2 phase relief)


Maximum pressure at spill-off valve discharge = 1001 kPa(a)

Key Results

Maximum pressure at spill-off valve discharge for this case = 1147 kPa(a)



Issues

34-023/1 Calculated maximum pressure at spill-off valve discharge exceeds value on control valve datasheet Rev C1

34-023/2 Is case where valve fails fully open considered?

See also 34-022/2




Calculation Book

34-024 / C MP Separator Max Relief Case - Network Analysis 02 24-Mar-93

Audit Tasks Methodology Consistency X See /1 As Built X See below

Key Assumptions


MP Separator Max Relief is gas blowby @ 249,332 kg/h (total load)

Based on 2 x 50% installed control valves CV = 450 each (not purchased at relief valve data sheet issue date)

PSV Datasheet at Rev C2 indicates multiple PSVs with staggered set pressures on MP Separator. 3 x 50% PSVs,

(2 operating + 1 standby). 2 are set at vessel DP + 1 set at 105% DP. Note that PSV datasheet on HOLD at Rev C2.

Relief is 100% vapour (no liquid or 2 phase relief)



Key Results




Issues


'As Built' P&ID shows 2 x 50% LVs (LV-7327/7332) + note stating max installed control valves CV = 350 each - therefore no

further action required. (Control valve datasheet states CV=330)



Calculation Book

34-025 / C 3rd Stage Compressor Max Relief Case - Network Analysis 01 27-Jan-93



3rd Stage Compressor blocked outlet relief is 134,720 kg/h for K-3303A and 134,720 kg/h for K-3303B. Relief from

one train at a time only considered

Staggered set pressures on 3rd Stage Compressor relief valves (3 x 33.3% PSVs 1st set at DP, 2nd at 103% DP & 3rd at 105% DP)


Maximum pressure at PSV discharge (Normal) = 1 to 500 kPa(g), (Built-up back pressure not considered)

Key Results




Issues

34-025/1 Calculated maximum pressure at PSV discharge exceeds value on PSV datasheet Rev C2 - Check for later revisions

34-025/2 Is relief from both compression trains a valid case?

HP flare isometric out of date when calc was made however unlikely to effect results significantly (line sizes

used are correct compared to 'As Built' P&ID).




Calculation Book

34-026 / C Injection Compressor Max Relief Case - Network Analysis 01 27-Jan-93

Audit Tasks Methodology Consistency X See /1 As Built Key Assumptions


Injection Compressor K-3304B blocked outlet relief is 155,273 kg/h

Relief from one train at a time only considered

Staggered set pressures on Injection Compressor relief valves (4 x 25% PSVs 1st set at DP, 2nd at 102% DP,

3rd at 104% DP & 4th at 105% DP)



Key Results




Issues


See also 34-025/2





Calculation Book

34-027 / C MP Separator Max Spill-off Case - Network Analysis 02 24-Mar-93

Audit Tasks Methodology X See /1 Consistency As Built Key Assumptions


MP Separator Max Spill-off (94,453 kg/h total load)




Key Results




Issues

34-027/1 Was failed open control valve considered?

See also 34-022/2




Calculation Book

34-028 / D West Test Separator Max Spill-off Case - Network Analysis 02 24-Mar-93



West Test Separator Max Spill-off (58,803 kg/h total load)




Key Results




Issues

34-028/1 Is case where valve fails fully open considered.

See also 34-022/2


used are correct compared to 'As Built' P&ID



Calculation Book

34-029 / D West Test Separator Max Relief Case - Network Analysis 02 28-Jan-93



West Test Separator max relief case is Blocked Outlet (189,279 kg/h total load). Equivalent to relief of full 30,000 bopd flowrate

1 x 100% PSV set at vessel DP

Relief is 2-phase


Maximum pressure at PSV discharge (Normal) = 1 to 500 kPa(g)

Key Results




Satisfactory 2-phase flow regime (annular)

Issues

None

See also 34-022/2




Calculation Book

34-030 / D East Test Separator Max Spill-off Case - Network Analysis 02 25-Mar-93



East Test Separator Max Spill-off (58,803 kg/h total load)




Key Results




Issues


See also 34-022/2



Calculation Book

34-031 / D East Test Separator Max Relief Case - Network Analysis 02 25-Mar-93



East Test Separator max relief case is Blocked Outlet (189,279 kg/h total load). Equivalent to relief of full 30,000 bopd flowrate

1 x 100% PSV set at vessel DP

Relief is 2-phase



Key Results




Satisfactory 2-phase flow regime (annular)

Issues

None

See also 34-022/2




Calculation Book

34-032 / D 1st Stage Compressor Max Relief Case - Network Analysis 01 28-Jan-93



1st Stage Compressor K-3301 blocked outlet relief is 72,801 kg/h

1 x 100% PSV set at 90% DP + 1 x 100% standby PSV set at 90% DP


Maximum pressure at PSV discharge (Normal) = 1 to 500 kPa(g)

Key Results




Issues

None

PSV set pressure is 90% downstream system design pressure to avoid tube rupture relief case on E-3302A/B





Calculation Book

34- / E 1st Stage Compressor Spill-off Case - Network Analysis 01 29-Jan-93

(No calc No.)



Max Spill-off = 59,841 kg/h (normal design flowrate)



Key Results




Issues

34-/1 Calculated maximum pressure at spill-off valve discharge exceeds value on control valve datasheet Rev C1

34-/2 Is case where valve fails fully open considered?



Calculation probably part of 34-032




Calculation Book

34-045 / E Total HP Blowdown Initial Conditions 01 22-Mar-93

(Checks blowdown line sizes for individual system blowdowns)



HP flare headers and sub-headers are not sized by blowdown case

Laterals are sized by individual section blowdown but sizing is velocity governed and not pressure governed

Blowdown simulations HIBBD155/156/109/110/111/112/164/114/115/104/105/185/184/106/113/165/107.LIS

Max velocity = 0.8 Mach

Individual blowdown lines 50m equivalent length

Considers pressure at end of lateral (i.e. header pressure) is atmospheric except for high flows where a header P is calculated

Compositions from blowdown simulations referenced above

ESI compressible flow analysis sufficient (network analysis using INPLANT not required)

Process data sheet for blowdown valves tagged on back of calc CM-E-C-K-M00-DS-0016 Rev D1, 11-Mar-93

Two types of blowdown valve:

- full bore ball valves (with downstream orifice installed)

- Angle type choke valves (or proprietary designed high pressure drop valves)

Key Results

Velocities in laterals found to be acceptable for given line sizes (i.e. less than Mach 0.8) except for lines from

'Dehydration' section (line size increased from 6" to 8") and 'Future Test Separator' section (line size increased from 3" to 6")

Issues

34-045/1 There is no network analysis run with common HP Blowdown at initial conditions

34-045/2 Consistency error in the blowdown flowrate from the gas injection flowlines (HP blowdown rates are identical to

those given in calc 34.006 except for GI flowlines where this calc uses 8 off BD valves at 2250 kg/h each and

calc 34.006 uses a combined figure of 8932 kg/h)

Other items to note are:

All line sizes are per 'As Built' P&IDs. Future 'Dehydration' section and 'Future Test Separator' section line sizes on latest P&IDs

(not 'As Built') have not been updated with results of this calc



Calculation Book

34-038 / F LP Separator Max Relief Case - Network Analysis 01 02-Mar-93


LP Flare system isometric No Rev Given

LP Separator Max Relief is gas blowby @ 110,874 kg/h (total load)

Based on 2 x 50% installed control valves CV = 700 each (not purchased at relief valve data sheet issue date)

PSV Datasheet at Rev C2 indicates multiple PSVs with staggered set pressures on LP Separator. 3 x 50% PSVs,

(1 operating + 1 standby are set at vessel DP + 1 operating set at 104.3% DP). Note that PSV datasheet on HOLD at Rev C2.


No tip P (pipe flare)


Key Results



Issues

None

Calculated maximum pressure at PSV discharge exceeds value on PSV datasheet Rev C2. However installed PSVs are

balanced so this will not affect the rating.

'As Built' P&ID shows 2 x 50% LVs (LV-7360/7367) + note stating max installed control valves CV = 650 each. Actual installed

valve CV=600 therefore no further action required.

Possible that LP flare isometric was out of date when calc was made however unlikely to effect results significantly (line sizes





Calculation Book

34-039 / F LP Separator Max Spill-off Case - Network Analysis 01 02-Mar-93



Max Spill-off = 60,147 kg/h



Key Results



Issues


See also 34-022/2





Calculation Book

34-040 / F Produced Water Surge Drum Gas Breakthrough Case - Network Analysis 01 02-Mar-93

(From LP Separator)



Produced Water Surge Drum Gas Breakthrough from LP Separator = 493 kg/h

Based on max CV of installed control valve (no datasheet in calc)



Key Results

Maximum Mach No. in system is negligible

Produced Water Surge Drum will not be overpressured for this case

Issues

None

Installed valve CV not stated in calc so impossible to check against P&ID however rate is so low as to be negligible for this study.


used are correct compared to 'As Built' P&ID)




Calculation Book

34-041 / F Produced Water Degassing Drum Gas Breakthrough Case - Network Analysis 01 02-Mar-93

(From MP Separator)

Audit Tasks Methodology Consistency As Built X See below

Key Assumptions


Produced Water Degassing Drum Gas Breakthrough from MP Separator = 54,938 kg/h

Based on max CV of installed control valve (no datasheet in calc)

Calc uses operating pressure of upstream vessel and max superimposed LP Flare backpressure of 142 kPa(a) downstream



Key Results


Produced Water Degassing Drum will not be overpressured for this case

Issues

None

Installed valve CV not stated in calc but resulting flowrate suggests a CV of 300 used. Actual installed CV = 330

which does not affect the results of the calculation significantly.





Calculation Book

34-042 / F Total LP Blowdown - Initial Conditions - Network Analysis 02 18-Mar-93


LP Flare system isometric Rev A3

Blowdown simulations HIBBD200/202/203/154.LIS

Total blowdown rate (initial rate) = 86,707 kg/h

Staggered flow - 3 minute time delay on Inj 'A' Compressor / Scrubber



Process data sheet for blowdown valves with calc CM-E-C-K-M00-DS-0016 Rev D1, 11-Mar-93

Key Results


Blowdown Valve Rev D1 Datasheet

backpressure

Network Analysis

Backpressure

33-ESV- kPa(a) kPa(a)

7279 201 266

7350 201 - No initial flow (staggered valve)

7372 201 231

7153 201 173

7177 201 203

7253 201 254

Additional valves on summary sheet not identified and no supporting blowdown valve datasheet attached

Issues

34-042/1 Total blowdown rate (initial rate) used in calc less than that in Relief & Blowdown Study Report ( 89,601 kg/h)

34-042/2 Validity of staggering blowdown. Were the systems sufficiently independent?

Some calculated backpressures greater than specified on blowdown valve Rev D1 datasheet. Not significant as still below

critical pressure ratio.






Calculation Book

34-043 / F Injection Compressor 'A' Blowdown - Initial Conditions - Network Analysis 02 18-Mar-93

Audit Tasks Methodology Consistency X See below As Built Key Assumptions


Blowdown simulations HIBBD201.LIS

Injection Compressor 'A' blowdown rate (initial rate) = 45,133 kg/h

No other equipment blows down at same time as Injection Compressor 'A'



Process data sheet for blowdown valves with calc CM-E-C-K-M00-DS-0016 Rev D1, 11-Mar-93

Key Results



backpressure

Network Analysis

Backpressure


7350 201 275

Issues

None

Calculation was prepared to identify the maximum velocity in the header (and consequently set the downstream pressure at

atmospheric for the worst case). Therefore network or back pressure results should not be used. The effect of the increase

in back pressure on the 'A' compressor blowdown valve is insignificant (no effect on critical pressure ratio).

See also 34-042/2





Calculation Book

34-033 / G Coalescer & LP Separator Heaters Simultaneous Fire Relief - Network Analysis 01 01-Feb-93



Considers pressure at end of sub-header (i.e. main header pressure) is atmospheric

Maximum allowable built-up back pressure at PSV discharge must be < 10% of set pressure (i.e. conventional valves)

PSV Relief Case Rate Configuration

31-PSV-

7378A/B Fire 40,771 kg/h 1 x 100% operation + 1 x 100% Standby



ESI compressible flow analysis sufficient

Key Results


Line sizes sufficient

Relief Valve Datasheet backpressure

Calculated Backpressure

31-PSV- kPa(a) kPa(a)

7378A/B 102-136 161

7428A/B 102-136 152

7437A/B 102-136 152

Issues

34-033/1 Assumption that the header is at zero pressure (i.e. that this is a singular event not coincident with any other releases)

Calculated backpressure greater than specified on datasheet - calc considers this OK as less than 10% of set pressure




Equipment Protected

LP Separator Heaters

Coalescer A

Coalescer B



Calculation Book

34-035 / G 3rd Stage Suction Scrubber PSV - Network Analysis 01 10-Feb-93


HP Flare system isometric Rev A1

Considers pressure at end of sub-header (i.e. main header pressure) is atmospheric (i.e. that this is a singular event not

coincident with any other releases)

Maximum allowable built-up back pressure at PSV discharge must be < 10% of set pressure i.e. conventional valves

- (suitability of conventional valves to be confirmed by inst / vendor - datasheet at rev C2, 11-Nov-92)

Relief Case Rate Configuration

Backflow 9,121 kg/h 1 x 100% operation


Key Results






7226 102-601 214.5

Issues

None

Basis for backflow calculation not given (probably NRV failure)

Possible that HP flare isometric was out of date when calc was made however unlikely to effect results significantly (line sizes




Calculation Book

34-036 / G Injection Stage Suction Scrubber PSV - Network Analysis 01 10-Feb-93





Maximum allowable built-up back pressure at PSV discharge must be < 10% of set pressure i.e. conventional valves - lots of margin

- (suitability of conventional valves to be confirmed by inst / vendor - datasheet at rev C2, 11-Nov-92)



PSV datasheet States 10% accumulation but 'Max Relieving Pressure' given is 121% of set pressure


Key Results






7326 102-601 174

Issues

34-036/1 Inconsistency on datasheet between accumulation and 'Max Relieving Pressure' (should be 10%)





PSV Equipment Protected

33-PSV-7326 Inj Stage Suction Scrubber B

33-PSV-7226 3rd Stage Suction Scrubber B




Calculation Book

34-037 / G HM & CM Expansion Drums Simultaneous Fire Relief Case - Network Analysis 02 01-Feb-93





Maximum allowable built-up back pressure at PSV discharge must be < 10% of set pressure (i.e. conventional valves)

Equipment Relief Case Rate Configuration

Protected

CM Exp'n Drum Fire 1,055 kg/h 1 x 100% operation + 1 x 100% Standby

HM Exp'n Drum Fire 27,628 kg/h 1 x 100% operation + 1 x 100% Standby


Key Results


Line size increase for common line from 6" to 8". Other line sizes sufficient

Datasheet backpressure

Calculated Backpressure*

kPa(a) kPa(a) * incorporating line size increase

102-136 136

102-136 185

Issues

34-037/1 Calculated backpressure (for 0152A/B) greater than specified on datasheet - calc considers

this OK as less than 10% of set pressure




PSV

64-PSV-0118A/B

63-PSV-0152A/B

64-PSV-0118A/B

63-PSV-0152A/B

Relief Valve



Calculation Book

34-044 / G Total LP Blowdown - After 3 mins (stagger point) - Network Analysis 01 22-Mar-93



Blowdown simulations HIBBD200/202/203/154.LIS

Total blowdown rate after 3 mins (stagger point) = 85,504 kg/h

Staggered flow - 3 minute time delay on Inj 'A' Compressor / Scrubber



Key Results

Calculation undertaken to check velocities were acceptable. Maximum Mach No. in system = 0.36 therefore line sizes sufficient


backpressure

Network Analysis

Backpressure


7279 201 176

7350 201 253

7372 201 157

7153 201 193

7177 201 157

7253 201 209

Additional valves on summary sheet not identified and no supporting blowdown valve datasheet attached

Issues

None

Total blowdown rate (initial rate) referenced in calc less than that in Relief & Blowdown Study Report ( 89,601 kg/h) but not

sufficient to effect sizing

Some calculated backpressures greater than specified on blowdown valve Rev D1 datasheet but not sufficient to affect sizing

See also 34-042/2





Calculation Book

34-046 / G Fuel Gas Cooler / Heater tube rupture relief line size check 01 02-Mar-93

Audit Tasks Methodology N/A Consistency N/A As Built X See /1

Issues

34-046/1 'As Built' P&IDs show bursting discs in this service (calc considers PSVs) therefore calc is no longer valid




Calculation Book

34-047 / G Simultaneous Fire Relief Case from E-6202 & Z-6201 A/B 01 02-Mar-93


HP Flare system isometric No Rev Given



62-PSV-0040A/B stated to be conventional valves yet 'Normal Back Pressure' is 1-500 kPa(g) (i.e. > 10% set pressure)

- others are balanced for similar 'Normal Back Pressure'


62-PSV-


0081 Fire 1,814 kg/h 1 x 100% operation

0092 Fire 1,814 kg/h 1 x 100% operation


Key Results


Header 3"-VH-34205 increased in size (shown as 3"-VH-34326 & 4"-VH-34327 on 'As Built' P&ID). Original size gave

mach no. = 0.97

Other line sizes sufficient


Calculated Backpressure*

62-PSV- kPa(a) kPa(a) * incorporating line size increase

0040A/B 102-601 153

0081 102-601 156

0092 102-601 160

Issues

None





Calculation Book

34-048 / G 2nd Stage Suction Scrubber A (D-3302A) PSV Discharge Line Size Confirmation 01 02-Mar-93





33-PSV-7099 balanced valve - suitability of balanced valve to be confirmed by inst / vendor - datasheet at rev C2, 11-Nov-92




Key Results






7099 102-601 157

Issues

None





Z-6201B


33-PSV-7099 D-3302A

Equipment Protected

E-6202 (Tube side)

Z-6201A



Calculation Book

34-049 / G Inj Stage Suction Scrubber A (D-3304A) PSV Discharge Line Size Confirmation 01 02-Mar-93





33-PSV-7301 conventional valve - suitability of conventional valve to be confirmed by inst / vendor - datasheet at rev C2, 11-Nov-92




Key Results






7301 102-601 188

Issues

None






Calculation Book

34-050 / G 3rd Stage Suction Scrubber A (D-3303A) PSV Discharge Line Size Confirmation 01 02-Mar-93


Key Assumptions








Key Results






7200 102-601 229

Issues

34-050/1 Rev C2 PSV datasheet states set pressure = 8200 kPa(g), 'As Built' P&ID shows set pressure = 7000 kPa(g)





33-PSV-7301 D-3304A



33-PSV-7200 D-3303A



Calculation Book

34-051 / G HP Fuel gas KO Drum (D-6201) PSV Discharge Line Size Confirmation 01 02-Mar-93



Considers pressure at end of PSV discharge line (i.e. sub-header pressure) is atmospheric (i.e. that this is a singular event not


62-PSV-0024A/B balanced valves


62-PSV-



Key Results






0024A/B 102-601 210

Issues

See 34-033/1





Calculation Book

34-052 / G E-3301 Shell Side PSV Discharge Line Size Confirmation 01 02-Mar-93







Fire 1,182 kg/h 1 x 100% operation


Key Results






0094 102-136 115

Issues

None




33-PSV-0094 E-3301


Equipment Protected

D-6201



Calculation Book

34-053 / G E-3303B Shell Side PSV Discharge Line Size Confirmation 01 02-Mar-93

Audit Tasks Methodology N/A Consistency N/A As Built X See below

Issues

None

'As Built' P&IDs show bursting discs installed in this service (calc considers PSVs) therefore calc is no longer valid

See also 34-046/1



Calculation Book

34-054 / G HP Manifold Relief - Network Analysis 01 02-Mar-93


Key Assumptions


Conventional valve

Calculation takes into account total system pressure drop but still assumes that this is a singular event not coincident

with any other releases


31-PSV-




Key Results







7042A/B 102-601 226

Issues

34-054/1 Rev C2 PSV datasheet states set pressure = 34,400 kPa(g), 'As Built' P&ID shows set pressure = 34,100 kPa(g)




Equipment Protected

HP Manifold



Calculation Book

34-055 / G Simultaneous Fire Relief Case from Z-3701 A/B, Z-3702 A/B & Z-6202 A/B 01 10-Feb-93



Balanced valves

Calculation takes into account total system pressure drop but still assumes that this is a singular event not coincident

with any other releases










Key Results







1021 102-601 344

1043 102-601 307

1063 102-601 336

1001 102-601 313

62-PSV-

0111 102-601 285

0122 102-601 282

Issues

See 34-033/1




37-PSV-1043

Z-3702A

Z-3701B

62-PSV-0122

Z-6202A

37-PSV-1063

Z-6202B

Z-3702B

62-PSV-0111

Z-3701A37-PSV-1001


37-PSV-1021



Calculation Book

34-056 / G Individual HP Separator Blowdown Case - Line Size Confirmation 01 10-Feb-93



Total blowdown rate (initial rate) = 41,311 kg/h



Key Results





backpressure



7318 551 402

Issues

None





Calculation Book

34-057 / G E-3701 Shell & Tube Side Simultaneous Fire Relief Case - Line Size Confirmation 01 12-Feb-93





Conventional valvesEquipment Protected Set Pressure Relief Case Rate Configuration

E-3701 (SS) 1380 kPag Fire 15,804 kg/h 1 x 100% operation

E-3701 (TS) 780 kPag Fire 2,135 kg/h 1 x 100% operation


Key Results


Calculation considers that line sizes sufficient




1482 102-136 197

1497 102-136 161

Issues

34-057/1 Calculated backpressure exceeds that specified on datasheet for both PSVs

See also 34-033/1




37-PSV-1497

PSV

37-PSV-1482



Calculation Book

34-058 / G E-6201A/B Tube Side Fire Relief Case - Line Size Confirmation 01 10-Feb-93





Conventional valves


62-PSV-

0001A/B Fire 828 kg/h 1 x 100% operation + 1 x 100% Standby


Key Results






0001A/B 102-601 112

Issues

See 34-033/1





Calculation Book

34-059 / G Comparative Program check of INPLANT Single Phase Simulation vs ESI 01 23-Apr-93


HP Separator Max Spill-off Case

Key Results

Pressure Drop given by ESI run ~20% less than INPLANT

Velocity given by ESI run 5% max less than INPLANT

Mach Nos. given by ESI run ~10% max less than INPLANT

Probably due to estimated average fluid properties used in ESI runs

Issues

34-059/1 Accuracy of calculations using ESI instead of INPLANT


E-6201A/B (Tube side)

Equipment Protected



Calculation Book

34-034 / G 2nd Stage Suction Scrubber B PSV - Network Analysis 02 26-Mar-93





33-PSV-7124 balanced valve - suitability of balanced valve to be confirmed by inst / vendor - datasheet at rev C2, 11-Nov-92




Key Results






7124 102-601 162

Issues

None






Calculation Book

34-064 / G HP Separator Max 2-Phase Relief - Network Analysis 01 07-Jun-93



HP Separator Blocked Outlet via PSV-7308A. Total load = 252,372 kg/h (based on 1 max well (30,000 bpd)

+ 1 average well (10,000 bpd) flowing)

Note at front of calc states that Rev 7 of Design Basis gives max well flow of 20,000 bpd + average well i.e. 30,000 bpd total

- calc considers 40,000 bpd flowrate anyway

INPLANT separator module not working therefore calc considers 2-phase flow to flare tip

No PSV datasheets included in calculation. Datasheet for PSV-7308 included in Calc 34-022 is for vapour relief only.

Tip P estimated from Kaldair supplied graph based on 64,273 kg/h vapour

Key Results


Line sizes appear sufficient




7308 102-851 646

Issues

None

See also PSV calculation technical audit


33-PSV-7124 D-3302B



Book H General

Book H contains a number of 'Check Print' copies of calculations reviewed in earlier volumes. There are nosignificant comments to record. In addition there a number of un-numbered calculation (all superseded). These 'Check Prints' and un-numbered calculations have not been reviewed in detail.


Calculation Book

34-001 / H Combined LP/HP KO Drum Sizing (Preliminary) No Rev 31-Jan-91

Audit Tasks Methodology N/A Consistency N/A As Built N/A

Calculation superseded. Not reviewed in detail



Calculation Book

34-002 / H HP KO Drum Sizing (Check) No Rev 04-Feb-91





Calculation Book

34-003 / H Flare Boom Length Calculations (Preliminary) No Rev 01-Mar-91





Calculation Book

34-004 / H Flare System - Material Balance for Flare UFD 0 25-Mar-91





Calculation Description Rev DateNumber

31.35 Relief Valve Calculations - HP Separator C1 Nov '91

Audit Tasks Methodology X see /1,4&5 Consistency X see /2&3 As Built X See below

Four Cases Considered; Fire, Blocked Outlet - Vapour Flow Only, Blocked Outlet - 2 Phase Flow (2 wells flowing),

Gas Blowby From Injection Compressor Suction Scrubber

Key Assumptions

Fire Case

Vessel dimensions: 3.7m x 18.8m T/T

No credit taken for vessel insulation

Blocked Outlet Case - Vapour Flow Only

Flow based 100% Hibernia normal case

Blocked Outlet Case - 2 Phase Flow (2 wells flowing)

2 wells flowing - 1 maximum well and 1 average well (total mass flowrate 207,848 kg/h)

Uses superseded API RP520 method of calculating separate orifice areas for liquid and vapour flow then adding together

Gas Blowby From Injection Compressor Suction Scrubber

Upstream pressure 138 bara

Valve CV = 16 max

Key Results


Case Relief Orifice Area

Flowrate Required

kg/h in2

Fire Case 38,835 1.55

Blocked Outlet Case - Vapour 227,649 9.38 Governing Case

Blocked Outlet Case - 2 Phase 207,848 3.19

Gas Blowby 34,192 Not Calc'd

Installed PSV orifice area from 'As Built' datasheet 9.6 in2.

Issues

31.35/1 Does 2 phase relief case become the governing case if the calculation new calculation method given

in API RP520, Seventh Edition used?

31.35/2 Flare network analysis for 2 phase case (Calc 34-064 / G) used total load = 252,372 kg/h (40,000 bpd).

31.35/3 Relief & Blowdown Study Report Rev C1 states HP Separator Blocked Outlet (Vapour) relief load is 244,897 kg/h.

31.35/4 The two phase calculation feed vapour / liquid split was abnormally low.

31.35/5 Methodological error in calculation (compared to API RP520 Sixth Edition). The wrong effective pressure was

for the V/L split and property conditions.

A note on the front of calc 34-064 / G states that Rev 7 of Design Basis gives max well flow of 20,000 bpd

+ average well i.e. 30,000 bpd total

HP Separator dimensions used in calc not 'As Built' (3.46m x 17.28m T/T) but this case is not governing therefore of no concern.

Actual installed control valve CV = 0.8 for blowby case.

Audited by AJR Date 07-Aug-00 Checked by MFG Date 16-Aug-00



31.36 Relief Valve Calculations - MP Separator C1 Nov '91

Audit Tasks Methodology X see /1-4,6-8 Consistency X see /5 As Built X See below

Six Cases Considered; Fire, Blocked Outlet - Vapour Flow Only, Blocked Outlet - 2 Phase Flow (2 wells flowing),

Gas Blowby From HP Separator, Gas Blowby From Test Separators (Individually), Gas Blowby From 3rd Stage Suction Scrubber

Key Assumptions

Fire Case



Blocked Outlet Case - Vapour Flow Only

Flow based 50% Hibernia, 50% Avalon case

Blocked Outlet Case - 2 Phase Flow (2 wells flowing)

2 wells flowing - 1 maximum well and 1 average well (50% Hibernia, 50% Avalon case)

Uses superseded API RP520 method of calculating separate orifice areas for liquid and vapour flow then adding together

Gas Blowby From HP Separator

Control valve upstream pressure 42.26 bara

Valve CV = 350 max (Revised CV in 1993. Originally 450)

Two LCVs in parallel installed but only one valve fails at any one time

Gas Blowby From Test Separator


Valve CV = 195 max

Gas Blowby From 3rd Stage Suction Scrubber


Valve CV = 16 max

Key Results



Flowrate Required

kg/h in2

Fire Case 27,241 2.30

Blocked Outlet Case - Vapour 75,227 8.50

Blocked Outlet Case - 2 Phase 88,590 2.22

Gas Blowby - HP Separator 249,332* 30.52* Governing Case

Gas Blowby - Test Separator 115,086 Not Calc'd

Gas Blowby - 3rd Stage Scrubber 10,107 Not Calc'd

* Based on original CV of 450

Installed PSV orifice area from 'As Built' datasheet = 2 x 16 in2 operating.

Issues

31.36/1 Does 2 phase relief case become the governing case if the calculation new calculation method given

in API RP520, Seventh Edition used?


31.36/3 Methodological error in calculation (compared to AIP RP520 Sixth Edition). The wrong effective pressure was

for the V/L split and property conditions.

31.36/4 The two phase calculation feed vapour / liquid split was abnormally low.

31.36/5 Calculation subsequently superseded but no indication that calculation was subsequently corrected.

31.36/6 The gas blowby cases are methodologically flawed.

31.36/7 The wrong control valve sizing equation is used in the calculation leading to an incorrect relief rate calculated

for the gas blowby from test separator case.

Actual installed HP Separator level control valve CV = 330 for blowby case.

Actual installed 3rd Stage Scrubber level control valve CV = 4.0.

MP Separator dimensions used in calc not 'As Built' (4.1m x 23.5m T/T) but this case is not governing therefore of no concern.




31.37 Relief Valve Calculations - LP Separator C0 27-Nov-91

Audit Tasks Methodology X see /1&2 Consistency As Built Four Cases Considered; Fire, Gas Blowby From MP Separator, Gas Blowby From Test Separators (Individually),

Gas Blowby From 2nd Stage Suction Scrubber

Key Assumptions

Fire Case



Gas Blowby From MP Separator


Valve CV = 700 max

Two LCVs in parallel installed but only one valve fails at any one time

Gas Blowby From Test Separator


Valve CV = 195 max

Gas Blowby From 2nd Stage Suction Scrubber


Valve CV = 16 max

Key Results



Flowrate Required

kg/h in2

Fire Case 14,195 3.28

Gas Blowby - MP Separator 110,874 28.80 Governing Case

Gas Blowby - Test Separator 24,853 Not Calc'd

Gas Blowby - 2nd Stage Scrubber 2,384 Not Calc'd

Installed PSV orifice area from 'As Built' datasheet = 2 x 16 in2 operating.

Issues

31.37/1 Is it possible for the Test Separator manifold to be connected to the LP Separator when operating in high pressure

mode?


See also 31.36/6

Actual installed MP Separator level control valve CV = 600 for blowby case.

Actual installed 2nd Stage Scrubber level control valve CV = 24.0. No concern as this is not a major relief case.




31.38 Inlet Line Size Checking for Relief Valves 05-Dec-91

Audit Tasks Methodology X see below Consistency X see /1 As Built X see /1

Key Assumptions

Inlet line equivalent length 100m

Preliminary data used for relief loads and selected orifice areas

Key Results

Relief valve inlet line sizes

Issues

31.38/1 Inlet line sizes should have been recalculated using 'Final' relief data and isometrics.

This calculation has obviously been revised as many PSV inlet line sizes are different to those calculated here.

The data used for this calculation is nearly all out of date. Some relief valve tag numbers have changed and all PSV relieving

capacities and maximum relieving capacities are different on the 'As Built' PSV datasheets. Some PSV set pressures are

also different.




31.41 Relief Valve Calculations - Oil Separation & Export C1 20-May-92

Audit Tasks Methodology Consistency See below As Built See belowEquipment Considered: E-3103 A/B LP Separator Heaters (HM Side)Key AssumptionsOnly fire case consideredExchanger dimensions: 1.75m OD x 7.431m OL (channel length = 1.1m)Wetted area calc takes into account 15m of 12" pipingPSV set pressure 1380 kPagShell is liquid fullLatent heat of vaporisation = 50 Btu/lbKey ResultsRelief flowrate = 40771 kg/h Required orifice area = 2.36 in2IssuesNone. 'As Built' exchanger has slightly different dimensions (1.175m OD x 7.641m OL) to those used in the calculation but considered insignificant.smaller diameter. Installed orifice area = 2.85 in2.Equipment Considered: D-3104 A/B CoalescersKey AssumptionsOnly fire case consideredCoalescer dimensions: 3.05m ID x 9.75 T/TWetted area calc takes based on LSHHPSV set pressure 700 kPagLatent heat of vaporisation = 415 Btu/lb (as per LP Separator)Key ResultsRelief flowrate = 7214 kg/h Required orifice area = 1.61 in2IssuesNone. Installed orifice area = 2.85 in2.Equipment Considered: E-3104 A/B Crude Product Coolers (HC & CM Side)Key AssumptionsOnly fire case consideredExchanger dimensions: 2.61m x 4.4m x 1.18mHC side PSV set pressure 700 kPagCM side PSV set pressure 1500 kPagWetted area calc for HC side uses 50% total surface areaWetted area calc for CM side uses 50% total surface areaLatent heat of vaporisation (HC) = 415 Btu/lb (as per LP Separator)Latent heat of vaporisation (CM) = 817.3 Btu/lb (treated as water)Key ResultsHC Side: Relief flowrate = 1827 kg/h Required orifice area = 0.41 in2CM Side: Relief flowrate = 928 kg/h Required orifice area = 0.15 in2IssuesNone. 'As Built' exchanger has slightly different dimensions (2.877m x 4.4m x 1.359m) to those used in the calculation but considered insignificant. Installed orifice area, HC side = 0.503 in2, CM side = 0.196 in2.Equipment Considered: E-3701 A/B Crude Recirculation Heater (HC & HM Side)Key AssumptionsOnly fire case consideredExchanger dimensions: 0.813m OD x 4.82m OL (channel length = 1.299m)Wetted area calc takes into account 20m for each sideShell is liquid fullHC side PSV set pressure 740 kPagHM side PSV set pressure 1460 kPagWetted area calc for HC (tube) side uses channel surface areaWetted area calc for HM (shell) side uses shell surface areaLatent heat of vaporisation (HC) = 182.1 Btu/lbLatent heat of vaporisation (HM) =50 Btu/lbKey ResultsHC Side: Relief flowrate = 4866 kg/h Required orifice area = 0.33 in2HM Side: Relief flowrate = 15804 kg/h Required orifice area = 1.13 in2IssuesNone. However, HC Side 'As Built' PSV datasheet has relief load = 2135 kg/hr and installed orifice area of 0.785 in2. This isequivalent to using the same properties in this calc as used for the Coalescer and LP Separator calculation (I.e. latent heat of vap. = 415 Btu/lb anf MW= 37.68).There is an error in the calculation as the orifice area req'd calc uses a flowrate in kg/h instead of lb/h. Resulting true requiredorifice area req'd should be = 0.74in2. As the installed orifice area is greater than the true required as given above there is no concern.Equipment Considered: Z-3701 A/B / Z-3702 A/B Crude Oil Pig Launcher / ReceiverKey AssumptionsOnly fire case consideredDimensions Launcher: 547.7 / 706mm ID x 7200mm OL

Receiver: 547.7 / 706mm ID x 11400mm OL Wetted area calc takes into account 15m of 12" pipingPSV set pressure 4500 kPagEquipment is liquid fullLatent heat of vaporisation = 50 Btu/lbKey ResultsLauncher: Relief flowrate = 12688 kg/h Required orifice area = 0.48 in2Receiver: Relief flowrate = 17345 kg/h Required orifice area = 0.66 in2IssuesNone. 'As Built' equipment has slightly different dimensions to those used in the calculation but considered insignificant.Installed orifice area = 0.785 in2 for both pieces of equiment.




31.43 Gas Blowby (Checking Capacity of Downstream System for Gas Blowby 22-Nov-92

from HP to MP Separator and MP to LP Separator)

Audit Tasks Methodology X see /1-3 Consistency See below As Built See below

Gas Blowby From HP to MP Separator

Key Assumptions

MP Separator PSVs orifice area is 2 x 16 in2 operating.

MP Separator spill off valve CV = 600

MP Separator spill off valve capacity during blowby case is 94,500 kg/h

Both 50% upstream LCVs fail at the same time

Key Results

Total relief capacity of PSVs and spill-off valve operating together is 365,594 kg/h

If HP Separator level control valves have a CV of 329.8 each, the total relief load if both valves fail open (365,594 kg/h) can

be handled by the installed PSVs and spill off valve operating together.

Issues

31.43/1 This calculation considers both upstream LCVs fail open simultaneously. This scenario is not considered in the

Relief & Blowdown Study Report Rev C1 (or in any other calculations reviewed), nor is the platform designed

for its affects.

Note that the 'As Built' HP Separator LCVs CV = 330 each , 'As Built' spill-off valve CV = 600 and 'As Built' MP Separator PSVs

orifice area is 2 x 16 in2 operating.

The maximum relief load given above (365,594 kg/h) is not considered in the Relief & Blowdown Study Report Rev C1 as

the governing HP flare relief case (HP Separator blocked outlet @ 244,897 kg/h) and is not considered in the flare hydraulic calculations

Gas Blowby From MP to LP Separator

Key Assumptions

LP Separator PSVs orifice area is 2 x 16 in2 operating.

LP Separator spill off valve CV = 4145

LP Separator spill off valve capacity during blowby case is 86,713 kg/h

Both 50% upstream LCVs fail at the same time

Key Results

Total relief capacity of PSVs and spill-off valve operating together is 194,875 kg/h

If MP Separator level control valves have a CV of 550 each, the total relief load if both valves fail open (194,875 kg/h) can

be handled by the installed PSVs and spill off valve operating together.

Issues

31.43/2 This calculation considers both upstream LCVs fail open simultaneously. This scenario is not considered in the

Relief & Blowdown Study Report Rev C1 (or in any other calculations reviewed), nor is the platform designed

for its affects.

31.43/3 The calculation identifies the failure of the spillover valve (open) could lead to a relief rate which is higher than

the current design.

Note that the 'As Built' LP Separator LCVs CV = 600 each , 'As Built' spill-off valve CV = 4145 and 'As Built' MP Separator PSVs

orifice area is 2 x 16 in2 operating. 'As Built' total relief load if calculated using the same method as given here will be higher

than 194,875 kg/h as the installed LCV CV = 600 (not 550).

The maximum relief load given above (194,875 kg/h) is not considered in the Relief & Blowdown Study Report Rev C1 as

the governing LP flare relief case (MP to LP Separator gas blowby @ 110,874 kg/h) and is not considered in the

flare hydraulic calculations


APPENDIX II

STAGE 2 PROPOSAL

Flare System Revalidation Study - Stage 2 Proposal Rev C (27 pages)

8266-HIB-TN-C-0001 Appendix II Revision: B/tt/file_convert/55276c7749795994178b46d5/document.doc Page 1 of 1 October 2000

Documents

Hibernia Study (Flare)