Pulp Digester failured

8/19/2019 Pulp Digester failured

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W W W. TA P P I . O R G JOURNALCORROSION 9Digester thinning:Erosion-corrosion of internalflow channel headersGreg A. Busby and Peter W. Hart

RECOVERY BOILER 19A laboratory study of recoveryboiler smelt shatteringAnton Taranenko, Markus Bussmann,and Honghi Tran

RECOVERY BOILER 27A novel method for determining the internal recycled dust loadin kraft recovery boilersMatheus Antunes Guimarães,Honghi Tran, and Marcelo Cardoso

CORROSION 37Could biomass-fueled boilersbe operated at higher steam temperatures? Part 1: Laboratory

evaluation of candidatesuperheater alloysDouglas L. Singbeil, Laurie Frederick,James R. Keiser, and W.B.A. (Sandy)Sharp

CORROSION 51Could biomass-fueled boilersbe operated at higher steam temperatures? Part 2: Field testsof candidate superheater alloysJames R. Keiser, W.B.A. (Sandy) Sharp,and Douglas L. Singbeil

CORROSION 65Could biomass-fueled boilers beoperated at higher steam temperatures? Part 3: Initialanalysis of costs and benefitsW.B.A. (Sandy) Sharp, W.J. (Jim)Frederick, James R. Keiser,and Douglas L. Singbeil

PULPING 81

Combustion properties of

reduced-lignin black liquorsNiklas Vähä-Savo, Nikolai Demartini,Rufus Ziesig, Per Tomani, HansTheliander, Erkki Välimäki, andMikko Hupa

T HE PAPER AND PACKAGING INDUSTRIES’ TECHNICAL RESOURC E AUGUST 2014 | VOL. 13 NO. 8

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MONTH 2007 |TAPPI JOURNAL 3

5 EDITORIAL Peer-Reviewed Engineering highlights from PEERS 2013 Peter W. Hart

9 CORROSION Digester thinning: Erosion-corrosion of internal flow

channel headers Greg A. Busby and Peter W. Hart

19 RECOVERY BOILER A laboratory study of recovery boiler smelt shattering Anton Taranenko, Markus Bussmann, and Honghi Tran

27 RECOVERY BOILER A novel method for determining the internal recycled dust

load in kraft recovery boilers Matheus Antunes Guimarães, Honghi Tran, and Marcelo Cardoso

37 CORROSION Could biomass-fueled boilers be operated at highersteam temperatures? Part 1: Laboratory evaluation ofcandidate superheater alloys

Douglas L. Singbeil, Laurie Frederick, James R. Keiser,and W.B.A. (Sandy) Sharp

51 CORROSION Could biomass-fueled boilers be operated at highersteam temperatures? Part 2: Field tests of candidatesuperheater alloys

James R. Keiser, W.B.A. (Sandy) Sharp, and Douglas L. Singbeil

65 CORROSION Could biomass-fueled boilers be operated at highersteam temperatures? Part 3: Initial analysis of costs

and benefits W.B.A. (Sandy) Sharp, W.J. (Jim) Frederick, James R. Keiser,and Douglas L. Singbeil

81 PULPING Combustion properties of reduced-lignin black liquors

Niklas Vähä-Savo, Nikolai Demartini, Rufus Ziesig, Per Tomani,Hans Theliander, Erkki Välimäki, and Mikko Hupa

WWW.TAPPI.ORG JOURNAL

TABLE OF CONTENTS VOL. 13 NO. 8

August 2014

» ON THE COVER: TAPPI’s Engineering Division promotes the application of engineering principles to

the design, construction, operation, and maintenance of facilities for the manufacture of pulp, paper,and related products. The Division’s Steam & Power/Energy Management and Corrosion & MaterialsEngineering Committees actively explore topics like the digester, recovery boiler, and biomass-fueledboiler research in this issue.

AUGUST 2014| VOL. 13 NO. 8| TAPPI JOURNAL 3

Editor-in-ChiefPeter W. Hart [email protected]+1 409 276-3465

Associate EditorScott Rosencrance [email protected]+1 770 429-2753

Editorial DirectorMonica Shaw [email protected]+1 770 367-9534

Vice President, Operations Eric Fletty [email protected]

PRESS Manager Jeff Wells [email protected]+1 770 209-7228

Webmaster Trina Heath [email protected]+1 770 209-7416

TJ EDITORIAL BOARDTerry L. Bliss, Ashland [email protected],+1 302 995-3523Brian N. Brogdon, FutureBridge Consulting & Training [email protected],+1 678 581-9114David A. Carlson, Carlson [email protected],+1 847 323-2685Jere W. Crouse, JWC [email protected],+1 608 362-4485Mahendra Doshi,Doshi and [email protected],+1 920 832-9101Peter W. Hart, MeadWestvaco [email protected],+1 409 276-3465Carl J. Houtman, USDA Forest Products [email protected],+ 1 608 231-9445John A. Neun, Albany [email protected],+1 518 275-8139Steven P. Ottone, SNP [email protected],+1 919 641-7854Arthur Ragauskas, Oak Ridge National [email protected]

John A. Roper III, Dow Chemical [email protected],+1 989 636-5644Scott Rosencrance, Kemira [email protected],+1 770 429-2753Ricardo B. Santos, MeadWestvaco [email protected],+1 540 969-2426Paul Wiegand, [email protected],+1 919 941-6417Junyong Zhu, USDA Forest Products [email protected],+1 608 231-9520

Join TAPPI today! TAPPI JOURNAL is a free benefit of TAPPI membership,and is only available to members. To join TAPPI, to renew yourTAPPI Membership,or to learn about other valuable benefits,visit www.tappi.org.TAPPI, 15 Technology Parkway S., Suite 115,Peachtree Corners, GA 30092, publishesTAPPI JOURNAL monthly.

ATTENTION PROSPECTIVE AUTHORS: All papers published are subject toTAPPI JOURNAL’s peer-review process. Not all papers accepted for reviewwill be published. Before submitting, check complete author guidelines athttp://www.tappi.org/s_tappi/doc.asp?CID=100&DID=552877.

Statements of fact and opinions expressed are those of individual authors.TAPPI assumes no responsibility for such statements and opinions. TAPPIdoes not intend such statements and opinions or construe them as asolicitation of or suggestion for any agreed-upon course of conduct orconcerted action of any sort.

Copyright 2014 by TAPPI, all rights reserved. For copyright permissionto photocopy pages from this publication for internal or personal use,contact Copyright Clearance Center, Inc. (CCC) via their website:www.copyright.com.If you have questions about the copyright permissionrequest process, please contact CCC by phone at +1 978 750-8400.To obtain copyright permission to use excerpts from this publication inanother published work, send your specific request in writing to Editor,TAPPI JOURNAL, 15 Technology Parkway S., Suite 115,Peachtree Corners, GA 30092, USA; or by fax to +1 770 4 46-6947.Send address changes to TAPPI, 15 Technology Parkway S., Suite 115,Peachtree Corners, GA 30092, USA; Telephone +1 770 4 46-1400,or FAX +1 770 446-6947. w ww.tappi.org

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The current issue of TAPPI Journal (TJ ) repre-sents the last in a series of special issues fea-turing highlights from the 2013 TAPPI PEERSConference held in Green Bay, WI, USA. The

rst two 2014 issues with peer-reviewed PEERS contentfocused on Pulp Manufacturing (March) and NonwoodPulping (June).

The papers in this issue were obtained from PEERSpaper sessions sponsored by the Engineering Division’sCorrosion & Materials Engineering and Steam & Pow-er/Energy Management Committees. Several committeechairs were asked to recommend papers appropriatefor an Engineering highlights issue of TJ. The authorsof the suggested manuscripts were solicited to submittheir work to the peer review process. The submissionand review process resulted in six papers available forthis special Engineering highlights issue.

A snapshot of Engineering issues The papers in this issue represent a snapshot of someof the ongoing research and development issues cur-rently being addressed by the Engineering Divisioncommittees in TAPPI. The topics include:

• Smelt — Continued effort is being expended toimprove the industry’s understanding of smeltshattering and smelt runoff events. Both of theseitems may be directly related to liquor quality andconsistency, as well as to the safe operation ofindustry equipment.

• Corrosion — Ongoing corrosion studies relate toequipment longevity and reduced maintenancecosts. Improved understanding of the type of andmechanisms associated with corrosion may alsolead to reduced capital costs for new pieces ofequipment.

Other Engineering topics at PEERS 2013 includedefforts in understanding the impact of new processingregimes, such as enhanced biomass boiler operation,that may improve capital requirements, maintenance,and efciency for large scale equipment. Also, theEngineering Division is actively following legislativeissues such as Boiler MACT and continuing to updatetheir constituents on the latest developments in bothproposed and currently active legislation that has thepotential to impact our industry.

An additional paper appears in this issue from

Niklas Vähä-Savo, a research associate and doctoralstudent at Åbo Akademi University in Finland. Thepaper on p. 81 is part of his Ph.D. thesis on “Behaviorof black liquor nitrogen in combustion – Formation ofcyanate,” to be defended later this year.

TAPPI PEERS 2014 in Tacoma The Engineering Division will again sponsor severalsessions at the 2014 TAPPI PEERS Conference in

Tacoma, WA, this September 14-17. The EngineeringProgram will consist of three distinct tracks: Corrosionand Materials; Energy, Power and Recovery; and

Editorial

Peer-Reviewed Engineeringhighlights from PEERS 2013


Editor’s Note: TAPPI would like to thank Brian Brogdon and Peter Hart of the TAPPI Journal (TJ) Editorial Board for soliciting papers from the 2013 PEERS Conference and overseeing their review for this special Engineering content, as well as additional past issues of TJ with PEERS content. Thanks also to the authors and peer reviewers who participated in this process.

– Monica Shaw, Editorial Director

PETER W. HART

Editor-in-chief


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Engineering and Maintenance. Additional tracks willbe sponsored by the Environmental Working Groupand the Pulp Manufacturing Division. Recyclingand Sustainability programs will also be offered. Inaddition, committee meetings and steering committee

meetings will be held during the conference. If youor your engineers want an excellent networkingopportunity, along with the opportunity to shape thecontent of future PEERS conferences, joining one ofthe Engineering Division’s subject matter committees(open to all TAPPI Members) provides an excellentstarting point. A list of committees is found at

www.tappi.org/Groups/Divisions/Engineering. We hope you enjoy the PEERS Engineering highlights

from TAPPI’s Engineering Division in this issue of TJ , and we hope to see you in Tacoma. TJ

Peter W. Hart, Ph.D., editor-in-chief of TAPPI Journal ,is the director of Pulping, Bleaching, and ChemicalProcess Technology for MeadWestvaco Corporation,Atlanta, GA, USA.

6 TAPPI JOURNAL | VOL. 13 NO. 8| AUGUST 2014

If you or your engineers want an excellentnetworking opportunity, along with theopportunity to shape the content offuture PEERS conferences, joining one of the Engineering Division’s subject matter

committees (open to all TAPPI Members)provides an excellent starting point.

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CORROSION


A s part of fitness for service and general safety guidelines,digesters are required to undergo routine shell thicknessinspections. The digester involved in this study was originallydesigned in 1993 as a 1250 a.d. tons/day, 2-vessel hydrauliccontinuous digester. Operation began in March, 1995. Thedesign consists of upper and lower internal flow channels forthe extraction and modified continuous cooking (MCC) zones

vs. external headers. The single-level of wash screens isequipped with an external ring header. The design is a Kamyr“wet feed system” consisting of a chip bin, chip meter, low-pressure feeder, steaming vessel, chip chute with a circulationpump, in-line drainers, and a sand separator in the circulationloop feeding the high pressure feeder. The digester has beenroutinely monitored and well maintained since it was firstcommissioned. Annual internal visual and ultrasonic thick-ness (UT) inspections through the years have always indicat-ed that the general corrosion rates were low. Thickness read-ings were typically taken at three locations on each shellcourse at eight sectors around the circumference. The typicalUT survey during the outages also included taking severalthickness readings of the shell that is directly accessible byremoving one of the flow channel clean-out windows. Usu-

ally, only one or two screens at each zone were opened for visual and UT of the shell, as well as wet fluorescent magnet-ic particle testing (WFMT) of shell welds or penetrant testing

(PT) of stainless nozzle welds. Since the internal collectionheaders were not equipped with drains that would removethe residual heel of colored extraction liquor below the outletnozzles, there was no visual inspection of the lower portionof the flow channels as viewed through the small clean-out

windows.In 2012, the mill asked for a six-month extension to the an-

nual inspection frequency. This necessitated a critical reviewof the available thickness data. During this review, it wasnoted that the UT data that had been taken inside the flowchannels through the clean-out windows varied by as muchas 0.300 inch around the circumference at the same digesterelevation. Even if the data were located on separate shellplates, all shell plates at any given elevation should have thesame nominal thickness. This variation in thickness was thefirst indication that the shell behind the flow channels hadexperienced in-service thinning.

Based upon review of the literature, rapid thinning can bea common problem in softwood continuous digesters runningmodern cooking processes. Wensley reported in 1998 thatcontinuous digester extraction liquors may produce corrosionrates up to nearly 0.25 in. per year [1]. Wensley’s paper also

showed that corrosion rates increase rapidly with increasingtemperature. A corrosion rate of 0.079 in. per year in a softwood continuous digester extraction liquor at 327°F increased

Digester thinning: Erosion-corrosion

of internal flow channel headers GREG A. BUSBY AN D PETER W. HART

PEER-REVIEWED

ABSTRACT: A softwood continuous digester has experienced severe erosion-corrosion of the shell wall insidethe internal flow channel headers of the extraction and modified continuous cooking (MCC) zones. Although erosion-corrosion is a common form, it is not typically looked for inside digester flow channels. The worst damage waslocated where high velocity liquor exits the screen orifices and enters the collection headers. With erosion-corrosionrates as high as 200 m/year, the damage has effectively reduced the wall thickness almost by half in the worst areas.Also affected were the horizontal backing rings that form the bottom of the flow channels. An API-579/ASME FFS-1Part 5, Level 2 analysis was performed to allow the mill to continue operating the digester until the next scheduledoutage. Owner-users are encouraged to inspect these locations in their digesters to ensure that erosion-corrosionhas not caused accelerated and/or unexpected shell thinning. The internal flow channels represent locations that arenot readily accessible for internal visual inspection without removal of the flow channel header cover plates andremoval of the residual black liquor typically retained in these flow channels. External, on-stream ultrasonicthickness testing (UT) at the proper collection header elevations has been determined to be an effective way todetect the presence of this erosion-corrosion phenomena. Also discussed are considerations to make on-stream UTas accurate as possible. Corrosion rates determined via electrochemical corrosion probe monitoring have correlatedvery well with the corrosion rates determined by on-stream UT data.

Application: The detailed results from this study should enable the users of continuous digesters with internalflow channels to improve their internal inspection techniques. Improved inspections will allow for safer operation ofthese type digesters.


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more than 3.5 fold to 0.290 in. per year when the temperatureincreased to 350°F.

Although equations are available to predict corrosion ratesin white liquor from the liquor composition, no predictiveequations have been developed to predict corrosion rates inextraction liquors. Predictive equations have not been corre-lated with the inorganic and organic constituents of extrac-tion liquor. The available equations do not account for theextraction liquor temperature, or the liquor flow velocity.

What is known is that extraction zone liquors are particularlycorrosive because of their high temperature, low caustic con-centration, and high flow rate.

PROBLEM IDENTIFICATION

The current work describes a softwood digester that has re-cently experienced severe erosion-corrosion of the carbonsteel shell behind the internal flow channel headers. For thesoftwood continuous digester in this case study, the combina-tion of liquor heaters operating wide open, a faulty alkali sen-sor that resulted in caustic concentrations below intended

values, and a high flow velocity inside the flow channels re-sulted in serious wall thinning damage to the shell. The dam-

age was greatest in the vicinity of the screen orifice holes, as well as the bottom 4 or 5 in. of the flow channels. Based uponobservations such as the erosion patterns and location of dam-

1. Roll-out reference sketch of digester.


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CORROSION


age, it was concluded the local flow velocity was a key param-eter affecting the rate of thinning damage. Descriptions ofhow the damage was initially found and how it was monitored

with on-stream ultrasonic thickness and electrochemical cor-rosion probe testing are reviewed. Also discussed are detailsspecifically related to external on-stream thickness testing,

which ensure that the corrosion rate data are accurate. A roll-out reference sketch of the digester in question, along withthe areas of high corrosion, is shown in Fig. 1 .

ONSET OF DIGESTER THINNING

The morphology of the damage inside the digester flow chan-nels clearly indicates the thinning is the result of erosion-cor-rosion. Erosion-corrosion is an acceleration of corrosive dam-age caused by the relative motion of a corrosive fluid againsta metal surface. When a certain velocity threshold is exceededfor a given alloy, the relative motion is capable of removing anotherwise protective (passive) film that forms on the metalsurface [2]. The thinning damage found in this digester ex-hibited the classic erosion-corrosion features such as grooves,

waves, comet-tails, teardrop-shaped pits, and horseshoe-shaped undercutting of the metal surface. Figure 2 showssome of the classic erosion-corrosion pictures obtained fromthe digester flow channels. Erosion-corrosion can producedramatic rates of damage, which were also experienced inthis digester. Based on the locations and patterns of thinning,the flow velocity appears to be the primary component of thedamage that occurred inside the digester’s flow channel head-ers. Static corrosion rates have always remained relatively nor-

mal for this digester.

RESPONSE TO DIGESTER THINNING

On-stream thickness testing (pre-outage) Once it was apparent that a hidden corrosion problem mayhave developed inside the internal flow channels, it was nec-essary to do external on-stream UT testing at the flow chan-nel elevations in order to verify the scope and magnitude ofdamage. Originally, small 4 in. x 4 in. windows were cut inthe insulation for thickness monitoring locations (TMLs).These TMLs were oriented 45° apart at the elevations thatcorresponded to the mid-height of the flow channel areas.The flow channels are actually 16 in. in vertical height, andthe true placement of the TMLs within the flow channel zone

was about midway. As the TML areas were scanned, the spot-to-spot thickness data varied considerably. The first impres-sion was that the UT readings were widely inaccurate; how-ever, it was determined that the true wall thickness actuallydid vary considerably from spot to spot in the TML. It wasalso clear that the thickness decreased near the bottom ofeach TML window. It was decided to remove the majority ofthe insulation corresponding to the flow channel zones. Thenew data collected indicated the true severity of thinning.

Extreme substrate loss was found near the bottom 5 or 6 in.of the flow channels and also in separate vertical bandsaround the circumference corresponding to the location of

the screen orifice holes. When plotted, the thinning had analmost sinusoidal pattern around the circumference of thedigester.

With the shell thickness as much as 0.300 in. below thecalculated ASME t-min [3] in the worst areas, the mill con-tracted with an engineering consultant to perform an API-579/

ASME FFS-1 fitness for service evaluation [4-5]. Using Part 5,a level 2 analysis was performed. Through the rigors of finiteelement analysis (FEA), it was possible to demonstrate that thedigester was indeed safe to operate for a few more monthsuntil the scheduled outage, which was months away.

On-stream ultrasonic thickness dataand accuracy

When trending UT data collected on a frequent basis, the dataneeds to be as precise and accurate as possible so that the indi-cated corrosion rates are not exaggerated. When ultrasonicthickness testing on operating digesters (on-stream inspection),three considerations are necessary for accurate and precisethickness data: 1) clearly identified and physically marked TML

locations directly on the digester surface; 2) temperature-veloc-ity compensation of the UT calibration; and 3) elimination oftemperature drift effects of the transducer as it heats up.

2. “Wavy” pattern of erosion-corrosion damage to the extractionheader plate directly beneath the screen orifice. In the lowerpicture, notice the erosion-corrosion damage inside the flowchannel in the vicinity of the extraction nozzle.

lower

extractionnozzle

weldcorrosion

shell


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CORROSION


As the temperature of a metal increases, the velocity ofsound through the material slows down; thus, it takes longerto go the same distance (thickness). This change in sound

velocity with temperature is an impor tant variable foraccurate readings. If the high temperature velocity effect isnot accounted for when testing a hot component, a UTtechnician will erroneously get a reading that indicates thecomponent is thicker than it actually is. For example, in the

wash zone of a typical continuous digester with a thicknessof 2 in. and an operating temperature of 320°F, the thicknesserror from operating temperature to ambient temperature(85°F) calibration will be about 0.050 in. [6]. ASME Section

V indicates a temperature correction factor of 1% per 100°Fdifference between the calibration and testing temperature.The alternate way for high temperature calibration is to heata thickness step block to the operating temperature and thenstore the calibration for this temperature corrected sound

velocity. Done with care, there should be excellent correlationbetween these two high temperature ultrasonic thicknesscalibration methods for the normalized SA 516-70 plate [7].

Another temperature induced error that can easi ly beeliminated is one related to high temperature transducerdrift. As the plastic potting material surrounding the trans-ducer heats up, the sound velocity passing through this mate-rial decreases even more than it would when passing througha metal. Since the measured sound must travel through thispart of the probe, as well through the digester shell, the tim-ing and gated thickness will be altered by this effect, referredto as “transducer drift.” An easy correction for transducer

drift is to set the thickness gage in the “echo-to-echo” mode.In this test mode, the machine is set to gate the thicknessbetween multiple echoes of the distance from the metal sur-face to the back wall. In this way, thickness readings in theecho-echo mode only represent the metal path distance andare not affected by the thickness of couplant, paint, or tem-perature drift. At typical digester operating temperatures andtesting techniques, transducer drift can create another 0.020to 0.030 in. of error in the thickness reading.

Internal inspection of digester flow channels— The moment of truth

Consistent with TAPPI Technical Information Paper TIP0402-27 “Guidelines for inspection of continuous digestersand impregnation vessels,” the internal inspection consistedof direct visual inspection, ultrasonic thickness testing, anddye penetrant and wet fluorescent inspection of critical

welds. Additionally, the f low channel cover plates were re-moved for the first time in the operating history of thedigester. Once the flow channel cover plates were removed,it was obvious that significant and very real damage had oc-curring within the flow channels. The damage was found tobe as bad as the on-stream UT data had indicated. Figures

3-5 are photos of the extraction flow channels showing theextent of corrosion that was present within the digester atthe time of the inspection.

3. Inside the lower extraction flow channel with cover removedfor access. Note the severe erosion-corrosion of the shellnear the bottom plate, as well as wash-out of the bottom plate

(holding liquor) that is directly beneath the screen orifice.

4. Inside the lower extraction flow channel with cover removedfor access.

5. Erosion-corrosion damage to the digester shell and bottom ofthe flow channel.

Bottom plate offlow channel withcover removed

Corrosion of bottomplate directly beneathscreen orifice

Digester shell

Flow channelbottom plate


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CORROSION


Extraction zoneThe upper and lower extraction zone consisted of a checker-board pattern of ten alternating carbon steel blank plates withten vertical stainless steel profile bar screens. The liquor col-lects through the screens and feeds through two flow orificesinto the flow channel header. The flow channel headers havetwo outlet nozzles oriented 180° apart. The erosion-corrosiondamage was concentrated near the flow orifice locations andaffected the lower portion of the shell inside the flow chan-nel. There was significant localized thinning of the bottomplate (backing bar) of the collection header where liquorflows through the screen orifices and impinges directly uponthe bottom plate (Fig. 5). At the lower extraction zone, severedamage extended 5 or 6 in. up from the bottom of the collec-tion header. Localized areas of shell thinning associated withthe “stand-off pins” were observed behind the screen platesas well as inside the flow channel. Galvanic corrosion effects

were not seen in these areas as the standoff pins do not actu-

ally touch the impacted area of the digester shell. Rather, thesepins tend to collect pulp fibers and increase the flow rate ofliquor around these pins. Figure 6 shows the localized cor-

rosion around the stand-off pins. Flow eddies appear to formaround these pins, which causes accelerated erosion-corro-sion with the classic “comet tail” and “horse-shoe track” pat-terns. The corrosion penetration into the shell at these stand-off pins located inside the flow channel were much deeperthan that around the pins behind the screens. Even thoughthe process chemistry in these two areas is nearly identical,higher velocities are present inside the flow channel than be-hind the screens, thus enhancing the erosion-corrosion rate.

MCC zoneThe MCC zone consists of two levels of diagonal, mill-slottedscreens with no blank plates. There are 20 screens in theupper zone and 20 screens in lower zone. At each screen,there is only one open flow orifice, as one of the two orificeholes was blanked in years past.

There was erosion-corrosion damage to the shell insideboth MCC flow channels, similar to that found at the extrac-tion zone. Although this damage required weld buildup repairto restore to ASME t-min thickness, the damage was much lesssevere than the damage at the extraction level. The bottomplates of the flow channel headers showed minor to moderatedamage underneath the open flow orifice. In general, thedamage to the shell in the MCC flow channels did not extendmore than 2 or 3 in. above the bottom plate of the collectionheader. Figure 7 shows the MCC channel along with the“heel” of liquor retained in the bottom of the flow channelthat impedes visual inspection of the internal flow channels.

WELD REPAIRS

All areas of the shell below the corresponding ASME t-min value were repaired by weld overlay. Prior to welding, the area was preheated to 300°F using a rose-bud technique [8]. The

welding was performed by shielded meta l arc welding(SMAW) using 5/32 dia., E7018 electrodes in consecutive,horizontal weld beads. The weld overlay areas were built-up

6. Note the localized damage around standoff pins that areattached to the flow channel covers that were removed. Thelower photo is a close up detailing the localized corrosioncaused by flow eddies around the standoff pins.

7. Upper MCC flow channel. (Nozzle height leaves 2.5 in. “heel”of liquor in the bottom).

erosion/corrosionentities inside lower

extraction flow channelcaused by flow eddies

around stainlessstand-off pins

blank plateremoved

(replaced)

screen (profilebars)

localized “wash-out”around stand-off pins

localized extraction flow channelwith covers removed


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CORROSION


to the corresponding t-min value plus an additional 0.050 in.of corrosion allowance. In a few of the worst areas of the lowerextraction zone, ER 309L SS electrodes were applied on topof the carbon steel weld build-up.

POST OUTAGE ON-STREAM

CORROSION MONITORINGBecause of the wide variance in local shell wall thicknessescaused by corrosion, TMLs used for UT comparison weremarked directly on the digester shell. Thickness readings

were taken at these precise locations for optimum corrosionrate monitoring accuracy. Due to the expected high erosion-corrosion rates inside the flow channels, the majority of TMLs

were concentrated at the flow channel elevations. Areasbelow the flow channels were included to establish “static”corrosion rates outside of flow areas.

Baseline thickness readings were taken prior to thedigester’s return to service. On-stream thickness readings

were taken approximately every 30 days for the next sixmonths. Annualized corrosion rates were calculated foreach TML to determine a short-term corrosion rate (betweenthe last two readings) and a longer-term corrosion rate(between the baseline and the last reading). These datashowed extremely high corrosion rates (100 to 200 m/year)in the majority of TMLs located in the flow channels and lowcorrosion rates (0-5 m/year) in static areas outside of theflow channels. Also, during the six months of on-streamthickness monitoring, the mill experimented with variouscooking schemes that showed a wide variance in corrosion

rates.In addition to the on-stream UT thickness monitoring, six

electrochemical corrosion monitoring probes installed by ananodic protection contractor during the recent outage werealso used to trend the corrosion rates during the six monthpost outage period. In general, there was very good correla-tion between the UT thickness data and the electrochemicalcorrosion rate data.

CONCLUSIONSOutside of the high velocity flow channel areas, the generalcorrosion rates on the digester shell range between 5 and 10mils/year. Inside the flow channels, directly below the screenorifice holes and along the bottom of the flow channels, thelocalized erosion-corrosion rate appears to be as high as 200mils/year. No thinning was noted behind the wash zonescreens, as there is an external ring header instead of an in-ternal flow header. Velocity appears to be the critical param-eter affecting the erosion-corrosion damage found inside theflow channels. Throughout the vessel, there has never beenany indication of stress corrosion cracking (SCC), and no in-dications of SCC were found during this outage. Because por-tions of the digester were well below the ASME t-min thick-ness, there was discussion about immediately shutting thedigester down for overlay weld repair. Instead, API 579/ASMEFFS-1 Fitness for Service Part 5 Level 2 analysis was performedto verify that the localized corrosion did not require an im-mediate shutdown of the digester. Rather, the digester wasallowed to run until the next scheduled outage. Weld repairs

were made to restore damaged areas to ASME t-min plus asmall corrosion allowance.

Based upon on-stream ultrasonic thickness data collected

since the outage, high erosion-corrosion rates still exist. How-ever, after experimenting with various cooking changes, the

ABOUT THE AUTHORS

We chose to study this aspect of digester thinning be-cause exploration of variations in routine thicknessinspections led to the surprising discovery of ex-treme rates of erosion-corrosion occurring within thedigester flow channels. The need to understand thecause, mitigation, and prevention of the rapid thin-ning led to the preparation of this work.

This work supports earlier efforts at identifyingvarious forms of rapid digester thinning experiencedin other continuous digesters. The current work com-pliments efforts in defining specific locations oferosion-corrosion.

Determining the root cause of the digester thin-ning and understanding the UT thickness testing datawere the main difficulties in this work. The most in-teresting thing we encountered was discovering theextent of digester thinning in areas that were deemedto be relatively safe and rarely inspected.

We have identified specific causes and locations ofpotential rapid digester thinning. Mills may use this

work to enhancetheir own di-gester inspec-tion programsto ensure safeoperation of thecontinuous di-gester. The nextstep includes re-building the lostbase metal andoverlaying thedigester flowchannels. Once this is done, routine monitoring ofthe flow channel thickness will be performed.

Busby is senior metallurgist for MWV Corp. inRichmond, VA, USA. Hart is director of Pulping,Bleaching, and Chemical Process Technology forMWV Corp. in Atlanta, GA, USA. Email Hart [email protected].

Hart Busby

http://www.etappijournal.com/tappijournal/august_2014/TrackLink.action?pageName=14&exitLink=mailto%3Apeter.hart%40mwv.comhttp://www.etappijournal.com/tappijournal/august_2014/TrackLink.action?pageName=14&exitLink=mailto%3Apeter.hart%40mwv.com


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CORROSION


mill determined the corrosion rate can be significantly var-ied. For instance, comparing UT survey data collected at thebeginning of down-flow cooking (where no extraction flowexists inside the extraction headers) with data 30 days afterthe cooking change showed the corrosion loss went from anannualized 0.190 in. per year to essentially zero. These chang-es in cooking conditions show that the liquor velocity is amajor component in the damage that has occurred withinthe flow channels. Being able to trial and error various cook-ing schemes and correlate these cooking schemes with cor-rosion rates only underscores the critical need for precise andaccurate on-stream ultrasonic thickness testing techniques.

For digesters with internal flow channels, if no currentor reliable wall thickness data from the shell near the bottomof the flow channel is available, owner/users are encouragedto schedule on-stream UT thickness measurements at theflow channel elevations or plan to remove flow channelcovers directly below screens at an upcoming outage forinspection. TJ

LITERATURE CITED

1. Wensley, A., “Corrosion of batch and continuous digesters,”Proc. -Int. Symp. Corros. Pulp Pap. Ind., 9th , CPPA, Montreal, 1998, p. 27.

2. Corrosion Technology Laboratory, “Erosion Corrosion, CorrosionFundamentals,” NASA. Available [Online] http://corrosion.ksc.nasa.gov/eroscor.htm>[31July2014].

3. ASME, “ASME Boiler and Pressure Vessel Code: An InternationalCode, 2013.” Available [Online] https://www.asme.org/getmedia/1adfc3df-7dab-44bf-a078-8b1c7d60bf0d/ASME_BPVC_2013-Brochure.aspx>[31July2014].

4. Antaki, G.,Fitness-for-Service Evaluations for Piping and PressureVessels: ASME Code Simplified , McGraw-Hill, New York, 2005.

5. Bennett, D.C. and Mixon, N., “API 579-1/ASME FFS-1, Fitness forservice post-construction code for pressure equipment with flawsand corrosion damage,”TAPPI PEERS Conf., TAPPI PRESS, Atlanta,GA, USA, 2011, p. 997.

6. Pezant, J.C., “High temperature thickness monitoring using ultra-sonic waves,” M.S. thesis, Georgia Institute of Technology, Atlanta,2008.

7. ASME Boiler and Pressure Vessel Committee on NondestructiveExamination, “ASME Boiler and Pressure Vessel Code, Section V:Nondestructive Examination,” American Society of MechanicalEngineers, New York, 1 July 2011. Available [Online]https://law.resource.org/pub/us/code/ibr/asme.bpvc.v.2010.pdf>[31July2014].

8. U.S. Department of Transportation Federal Highway Administration,“Guide for Heat-Straightening of Damaged Steel Bridge Members,”Chap. 2, Section 2.7. Available [Online] http://www.fhwa.dot.gov/BRIDGE/steel/02.cfm >[31July2014].

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I n kraft recovery boiler operations, molten smelt (mainlyconsisting of sodium carbonate and sodium sulfide, andsmall amounts of sodium sulfate, sodium chloride, andpotassium salts) flows out of the boiler at 800°C-850°Cdown a number of spouts, at a flow rate on the order of1 L/s per spout, before falling into the dissolving tank below

( Fig. 1 ). In the tank, the hot smelt mixes with weak wash toproduce green liquor that is subsequently causticized withlime to produce white liquor for reuse in the pulping process.The interaction between the molten smelt and water in theconfined space of the dissolving tank is violent; the noise and

vibration of the tank often can be heard and felt far from thetank itself. To control the intensity of the smelt-water interac-tion, mills use steam “shatter jets” to break up the smelt streaminto a spray of droplets just below the ends of the spouts.

Intense smelt-water interaction may be necessary to effec-tively dissolve smelt in the dissolving tank. Smelt shattering

is important to distribute smelt evenly throughout the tank,rather than have large amounts of smelt simply pour into thetank from the spout. Inadequate smelt shattering increasesthe violence of dissolving tank smelt-water interaction. Expe-rienced boiler operators claim to be able to assess dissolvingtank operation by listening to the tank. At the extreme, inad-equate smelt shattering can lead to a dissolving tank explo-sion that can cause equipment damage, an unscheduled shut-down, and even personnel injury [1,2]. Despite theseconcerns, smelt shattering practices vary widely from mill tomill, and the shattering behavior has not been studied before.The safety implications and lack of standards for smelt shat-tering motivated the study presented here.

CURRENT PRACTICE

An informal survey of a number of mills led us to concludethat smelt shattering practices vary widely. Shatter jet nozzlescome in a variety of designs. Figure 2 shows examples of afew designs that we encountered during mill visits. Somemills simply modify the end of a steam tube to create a singlenozzle (Fig. 2a), or attach a T to create multiple nozzles(Fig. 2d). Shatter jet nozzles can be round (Figs. 2a, 2b, and2d) or slit-shaped (Figs. 2c and 2e). Frozen smelt can build up

on shatter jet nozzles, which has led some mills to install thenozzle within a guard (Fig. 2b) or to use hot water or weak wash to clean the shatter jet nozzle (Fig. 2a).

A laboratory study of recovery

boiler smelt shattering ANTON TARANENKO, MARKUS BUSSMANN, AN D HONGHI TRAN

RECOVERY BOILERPEER-REVIEWED

ABSTRACT: A scaled-down experimental apparatus was built to examine smelt shattering during typical recov-ery boiler operations. Water-glycerine solutions and air were used in place of smelt and steam. A high-speed cameraand image processing software were used to record and quantify liquid shattering in terms of droplet number andsize distributions, as a function of air velocity, air nozzle position, liquid flow rate, and liquid viscosity. The resultsshowed that increasing shatter jet velocity reduced average droplet size, increasing the liquid flow rate increaseddroplet size, and placing the shatter jet nozzle closer to the liquid stream decreased droplet size. These results wereall as expected. The effect of liquid viscosity (1-50 cP) depended on the shatter jet velocity. At high air velocities,even the viscous liquid was well shattered, but at lower velocities, the effect of viscosity on shattering wassignificant. Application: Understanding how a molten smelt stream is shattered by a steam jet will help operators and pro-cess engineers to optimize their recovery boiler smelt shattering efficiency and to improve dissolving tank safety.

1. Typical smelt shattering practice.


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Steam used for smelt shattering varies from mill to mill, buttypically measures 3-15 bar (45-220 psi) and 150°C-250°C.Steam consumption also varies widely, from 180 kg/h to 2250kg/h per nozzle, or an order of magnitude range.

Shatter jet nozzle placement and orientation and spout in-clination can also affect shattering. Shatter jets are usually

aimed at the smelt flow just below the end of the spout, andthe nozzles are usually installed above the spout and pointdownwards, or further from the boiler wall and point backtowards the smelt flow. The choice affects how well smelt canbe shattered when the flow rate is unusually high or low,

which along with spout inclination, affects the trajectory ofthe smelt. When the smelt flow rate is low, and especially ifthe spout inclination is shallow, smelt will simply drip off thetip of the spout. When the smelt flow rate is high, and espe-

cially if the spout inclination is steep, smelt will shoot off thespout. In either case, shatter jet placement will affect shatter-ing effectiveness. To accommodate such situations, some millsinstall two jets per spout and turn on both during periods ofhigh smelt flow. Some mills also install adjustable shatter jetsthat an operator can point toward abnormal smelt flows.

EXPERIMENTAL SETUP AND METHODOLOGY

A laboratory-scale experimental setup ( Fig. 3a ) was con-structed to study smelt shattering. A water-glycerine solution(liquid) was used in place of molten smelt, and compressedair in place of steam. Liquid shattering was characterized asa function of air velocity, nozzle position, liquid flow rate, andliquid viscosity by imaging the spray with a high speed cam-era (Fig. 3b), and then processing those images to extract

3. Shattering apparatus: a) schematic and b) illustration.

2. Examples of shatter jet nozzle designs.


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droplet number and size distributions. Although the appara-tus was built at a reduced scale, the various parameters werechosen to reflect typical values of these parameters in recov-ery boilers.

As shown in Fig. 3a, the apparatus was operated as follows. A large collection tank stored the water-glycerine solution.The solution was pumped through a flow meter to an inclinedtank, with a spout mounted near the top (Fig. 3b). After thattank filled, the liquid began to flow down the spout (the in-clination was maintained at 15° for all of the experiments de-scribed here) at a known flow rate, and fell back into the col-lection tank.

The liquid stream was shattered into droplets by an imping-ing air jet (Fig. 3b). The air was drawn from the building sup-ply. The air flow rate was controlled by a ball valve and mea-sured by a flow meter. The air line was capped with aLaval-type nozzle (11.9 mm outlet dia., 7.9 mm throat dia.).The nozzle was positioned either 7.5 cm or 15 cm above theliquid stream and pointed straight down. This nozzle proxim-ity to the liquid stream, if scaled up, translates into 30-60 cm,a typical distance between a shatter jet nozzle and the smeltstream.

Shattering experiments were conducted at liquid f lowrates of 0.1 L/s and 0.2 L/s to yield liquid velocities at the endof the spout similar to those in practice. The lower flow rate(0.1 L/s) is representative of the smelt flow encountered undernormal recovery boiler conditions; 0.2 L/s is more representa-tive of heavy smelt flow. Different water-glycerine solutions

were used to obtain liquid viscosities of 1, 2.5, 10, and 50 cP,

where 3-5 cP is typical of recovery boiler smelt at 800°C-850°C[3]. (Liquid viscosity was limited to 50 cP because more vis-cous liquids could not be pumped at the requisite 0.2 L/s.) Theair velocity at the shatter jet nozzle exit was set to 100, 150,200, 250, or 300 m/s, corresponding to inline air pressures of10, 12.5, 15.5, 18.5, and 21.5 psig.

Image analysis

The liquid spray was backlit by a 300 watt light source andimaged by a Mega Speed MS70K S2 digital high speed camera(Canadian Photonics Lab, Inc.; Minnedosa, MB, Canada) fitted

with a 28-300 mm f/3.5-6.3 DG macro lens (Sigma Corp.;Kanagawa, Japan). A translucent sheet behind the spray wasused to diffuse the light, which then shone through the sprayand into the camera. Figure 4 shows a characteristic spraypattern from the end of the spout to 50 cm below, and theposition of the 3.5 × 3.5 cm section of the spray that was im-aged. The depth of field of the images was about 2 cm.

After a set of images of a particular spray configuration hadbeen obtained, open source software (ImageJ) was used toprocess the images and extract droplet size data. Figure 5 depicts the processing of a sample image. The original image(Fig. 5a) was first converted into a binary (black and white)

format (Fig. 5b). ImageJ was then used to calculate individualdroplet areas (in pixels) from the binary image, which werethen converted into equivalent droplet diameters. Only circu-

4. A view of the overall spray, and of the section that wasimaged (not to scale).

5. Image processing: a) original image, b) binary of the originalimage, and c) droplet outlines.


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lar droplets were analyzed; larger liquid fragments, and drop-lets that were out of focus, were discounted. Figure 5c showsthe droplets that ImageJ identified from Fig. 5a.

Data from multiple images of any spray configuration wasthen consolidated into droplet number and size distributions,and the Sauter mean diameter (SMD or D 32 ) was calculated.The SMD is a common measure of the fineness of a spray; itreflects the average ratio of drop volume to surface area [4,5]:

(1)

where D is the droplet diameter and N is the number of droplets.

RESULTS AND DISCUSSION

Results are presented of the effect of four parameters on liquidshattering: air velocity (u air ), liquid flow rate (Q l ), liquid viscos-ity ( µ l ), and the distance between the nozzle and liquid (N ls ).Images were processed for each of 80 experimental condi-tions, to allow for an analysis of the effect of each of the pa-rameters while keeping the others constant.

Effect of air velocity We considered five air flow rates, characterized by averagenozzle exit velocities of 100, 150, 200, 250, and 300 m/s, as

measured by the air flow meter. Axial velocities downstreamof the nozzle were then measured with a pitot tube to assessthe rate at which the air jets decayed. Velocities were mea-sured horizontally, vertically, and diagonally across the jets.The results confirmed that the jets were approximately axi-symmetric. Figure 6 presents a sample measurement ofhorizontal air velocity profiles measured 7.5 cm from the noz-zle exit. The nozzle exit diameter was 11.9 mm. Figure 6 il-lustrates jet profiles six nozzle diameters downstream of theexit. The jet center line velocities are only a half to a third ofthe average nozzle exit velocity, and the jet is several times as

wide as at the nozzle exit.Figure 7 displays a set of representative images that il-

6. Shatter jet air velocity distribution; N ls = 7.5 cm.

7. Spray images depicting the effect of air velocity on shattering; Q l = 0.1 L/s, μL = 2.5 cP, N ls = 7.5 cm.


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lustrate the effect of air velocity on liquid shattering. In thiscase, Q l = 0.1 L/s, µ l = 2.5 cP, and N ls = 7.5 cm. As expected,the stronger the jet, the smaller the droplets, as a result of anincrease in the aerodynamic drag applied to the liquid stream.The 100 m/s jet is not strong enough to adequately shatterthe liquid stream, as the image shows several large drops ofliquid. Increasing the velocity of the shatter jet by 50 m/slargely eliminates those large drops, and further incrementsof air velocity lead to a progressively finer spray.

Figure 8 shows two droplet size distributions as a func-

tion of air velocity. The number distribution (Fig. 8a) can bemisleading, as it appears to indicate that shattering is nearlyindependent of air velocity. The problem is that even inade-

quate shattering yields many small droplets, but also a fewlarge drops that contain a large fraction of the total liquid

volume. The liquid volume distribution (Fig. 8b) conveys abetter sense of the effect of air velocity. The plot shows thatincreasing air velocity shatters more of the liquid volume intosmall droplets, and that large drops only appear at low air

velocities.Figure 9 shows SMD versus air velocity and quantifies

the extent to which mean droplet size decreases with air ve-locity. In this case, the mean droplet size decreases by half, asthe air flow rate increases three-fold.

Effect of liquid flow rate To examine the relationship between liquid flow rate andaverage liquid velocity, the liquid cross-sectional area wasimaged at the end of the spout ( Fig. 10 ) for various f lowrates and spout inclinations (0°, 10°, 15°, and 20°) and usedto calculate an average liquid velocity that is plotted inFig. 11 . One can draw two conclusions from this plot. At a

given inclination, the velocity is almost independent of flowrate. This means that the water level in the spout rises withincreasing flow rate. As long as the spout is inclined, the ve-

8. The effect of velocity on droplet size distribution: a) numberdensity and b) volume density; Q l = 0.1 L/s, μL = 2.5 cP, N ls =7.5 cm.

9. The effect of air velocity on Sauter mean diameter; Q l =

0.1 L/s, μL = 2.5 cP, N ls = 7.5 cm.

10. Liquid flow down a spout (front view).

11. Effect of liquid flow rate on spout exit velocity at differentinclination angles.


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locity is not a strong function of the inclination. It is only inthe horizontal spout that the liquid flows more slowly, andthus the water level is higher.

To examine the effect of liquid f low rate on shattering, twoflow rates were considered: 0.1 L/s and 0.2 L/s. Figure 12 shows the SMD as a function of air velocity for each of the twoliquid flow rates. For these particular conditions, doubling theliquid flow rate roughly increased the mean droplet diameterby almost a factor of two.

Effect of liquid viscosity The effects of liquid viscosity on the flow down the spoutand on shattering were studied. Figure 13 shows the aver-age liquid velocity at the spout exit versus flow rate for fourdifferent viscosities. As expected, as viscosity increased, theliquid velocity decreased (and the liquid level in the spout

increased).Figure 14 shows the effect of liquid viscosity on shatter-ing. The 300 m/s air jet shattered all of the liquids well, as

demonstrated by the small variation in mean droplet diameter.But when the air velocity was only 100 m/s, the effect of vis-cosity was significant; the SMD varied from 1.2 mm for the 1cP liquid to 2.1 mm for the 50 cP liquid. Liquid viscosity obvi-ously matters when the shatter jet is not strong enough tosimply overpower the liquid stream.

Effect of nozzle proximity Figure 15 shows the effect of nozzle position on droplet size(Q l = 0.1 L/s and µ l = 2.5 cP). As expected, placing the nozzlecloser to the liquid stream results in improved shattering.Droplet mean diameter decreased as the shatter jet nozzle waspositioned closer to the liquid stream. Roughly speaking, theSMD decreased by half when the nozzle is moved from 15 cmto 7.5 cm from the liquid, for all air velocities.

PRACTICAL IMPLICATIONS

The results of this laboratory-scale study confirm what is

14. The effect of liquid viscosity on Sauter mean diameter;Q l = 0.1 L/s, N ls = 7.5 cm.

12. The effect of liquid flow rate on Sauter mean diameter;μ l = 2.5 cP, N ls = 7.5 cm.

15. The effect of nozzle proximity on Sauter mean diameter;Q l = 0.1 L/s, μ l = 2.5 cP.

13. Liquid velocity at the spout exit versus viscosity; spoutinclination = 15°.


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common sense: shattering improves with increased shatter jet flow rate, it improves as the shatter jet nozzle is movedcloser to the liquid stream, and it is easier to shatter the liquidat lower flow rates than higher ones, as long as the nozzle isproperly positioned. For the range of liquid viscosities that

we studied (1-50 cP), the results show that adequate shatter-ing requires a minimum shatter jet flow rate that must in-crease with liquid viscosity.

Although our results are prel iminary, we can begin todevelop correlations that can be applied to actual scale. Acommon parameter that is often used to characterize liquidatomization (shattering) is the liquid-to-gas momentum ratioq [6]:

(2)

where ρ l and ρ g are the liquid and gas densities, and u l and u g are the liquid and gas velocities. Figure 16 shows mediandrop diameter normalized by the diameter of the liquidstream at the end of the spout, D l, plotted against themomentum ratio q .

As expected, as the liquid-to-gas momentum ratio decreas-es, the mean drop size decreases as well. Despite considerablescatter in the data, the results are generally consistent for thetwo liquid flow rates that we considered. This is an exampleof a result that we plan to build on to generalize shatteringbehavior, and ultimately to reliably scale up the results.

Additional experiments are planned to examine the effec-tiveness of different nozzle geometries and positions, and to

examine an even larger range of flow rates and viscositiesthat might be encountered during boiler startup and upsetconditions. TJ

ACKNOWLEDGEMENTS

This work was conducted as part of the research program,“Increasing energy and chemical recovery efficiency in thekraft process – III,” jointly supported by the Natural Sciencesand Engineering Research Council of Canada (NSERC) and aconsortium of the following companies: Andritz, AV Nacka-

wic, Babcock & Wilcox, Boise, Carter Holt Harvey, Celulose

Nipo-Brasileira, Clyde-Bergemann, DMI Peace River Pulp,Eldorado, ERCO Worldwide, Fibria, FP Innovations, Interna-

16. Liquid-to-gas momentum ratio versus normalized dropletmean diameter.

ABOUT THE AUTHORS

The way in which molten smelt is shattered by asteam jet is believed to greatly affect dissolving tankoperation and safety. We chose this topic to study toprovide better insight into the shattering behavior ofmolten smelt.

To our knowledge, no previous research has beendone on this topic. The most difficult aspect of this re-

search was the scarcity or non-existence of literatureon smelt shattering. We addressed the problem byvisiting a number of pulp mills, consulting with millengineers and operators, and analyzing smelt shatter-ing videos.

It was surprising to find that increasing the liquidflow rate in an open channel did not increase its ve-locity, and that increasing the liquid viscosity from 1cP to 50 cP had little ef fect on the liquid droplet size.

Understanding how a molten smelt stream is shat-tered by a steam jet might help mills to optimize theirsmelt shattering efficiency and to improve their dis-solving tank safety.

The next step is to continue to use the laboratory

apparatus to examine the effects of different jet noz-zle designs and positions on droplet size anddistribution.

Taranenko is graduate student, Tran is professor,Pulp and Paper Centre, and Department of ChemicalEngineering & Applied Chemistry, and Bussmannis associate professor, Department of Mechanical &Industrial Engineering, University of Toronto, Toronto,

ON, Canada. Email Tran at [email protected].

Trann Bussmann Taranenko

http://www.etappijournal.com/tappijournal/august_2014/TrackLink.action?pageName=25&exitLink=mailto%3Ahonghi.tran%40utoronto.cahttp://www.etappijournal.com/tappijournal/august_2014/TrackLink.action?pageName=25&exitLink=mailto%3Ahonghi.tran%40utoronto.ca


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tional Paper, Irving Pulp & Paper, Kiln Flame Systems, Klabin,MeadWestvaco, Metso Power, StoraEnso Research, Suzano,Tembec, and Tolko Industries.

LITERATURE CITED

1. Grace, T.M. and Tran, H.N., “Critical issues in dissolving tank oper-ation,” Int. Chem. Recovery Conf., TAPPI PRESS, Atlanta, GA, USA,2010.

2. Lien, S. and DeMartini, N., “Dissolving tank explosions: a reviewof incidents between 1973 and 2008,” unpublished report, BlackLiquor Recovery Boiler Advisory Committee and American Forest &Paper Association, New York, 2008.

3. Tran, H.N., Sunil, A., and Jones, A.K.,J. Pulp Pap. Sci. 32(3):182(2006).

4. Yule, A.J. and Dunkley, J.J., inAtomization of Melts for PowderProduction and Spray Deposition , Clarendon Press, Oxford, NY, USA,1994, Chap. 3, pp. 47-87.

5. Lefebvre, A.H., inAtomization and Sprays , Hemisphere Publishing,New York, 1989, Chap. 3.

6. Mashayek A. and Ashgriz, N., inHandbook of Atomization and Sprays:Theory and Applications (N. Ashgriz, Ed.), Springer, New York, 2011,Chap. 29, pp. 657-684.

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I n kraft recovery boilers, the hot flue gas from black liquorcombustion gives off heat as it passes through a series ofheat transfer tube surfaces in the upper boiler to producesteam. The cooling causes the vapors of alkali compounds inthe flue gas to condense and form fumes ( Fig. 1 ). This, to-gether with a small amount of physically entrained carryover,forms a stream of fly ash or dust that is collected by ash hop-pers and electrostatic precipitators (ESP) [1]. A small portionof the dust stream ends up forming deposits on the tube sur-faces, which are occasionally knocked off by sootblowers andfall to the char bed or to ash hoppers. Only a small amount(


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of fume formed thermodynamically and equating the amount with the IRD load, and (3) calibrating the calculated dust loadagainst the actual dust load determined independentlythrough a material balance around the mix tank over a4-month period. The study showed that the increasing charbed temperatures increased the internal IRD load, and that atthe same average bed temperature, a bed with a poor tem-perature distribution generated twice as much IRD as a bed

with a good temperature distribution.In another study, Tamminen et al. [6,7] developed an on-

line recovery boiler dust analyzer and used it to measure theamount and composition of dust in the flue gas in two recov-ery boilers. The analyzer operated by continuously collectingthe flue gas dust sample, dissolving it in water, and analyzingthe flowing water potentiometrically for ion concentrations.They found that at full liquor firing loads, the average dustloads in these two recovery boilers were 6 wt% and 7 wt% as-fired BLDS, respectively. Their data showed that the amountof dust increased with an increase in liquor firing load and infurnace floor surface area, and that the majority of the fume(90%–95%) was formed from in-flight burning of black liquordroplets; only 5%–10% came from the char bed.

As more and more kraft pulp mills have adopted high-sol-ids firing practices in recent years, many recovery boilers op-erate at a high bed temperature. This inevitably results in anincrease in fume formation, and hence in IRD load. While

wide variations in IRD load can have a great impact on the as-fired black liquor mass flow, composition, heating value,steam production, and ultimately on the boiler thermal per-

formance, there is no simple and reliable means of directlymeasuring this dust stream. As a result, not many mills knowhow large a recycled dust stream is circulating around theirrecovery boilers.

In this paper, we first review the principles of several meth-ods that have been used to estimate the amount of IRD. Wethen discuss the results of a systematic study conducted onrecovery boilers at three kraft pulp mills within Fibria Celu-lose SA. The study was part of a larger effort to develop asimple method that could be used to reliably determine IRDload in recovery boilers, and then use it to investigate keyboiler operating variables that lead to high IRD dust loads.

COMMON METHODS

Mass balance method As-fired black liquor is the virgin black liquor that has beenmixed with dust collected from ESPs and ash hoppers. There-fore, the amount of IRD is essentially the mass difference be-tween the virgin black liquor and the as-fired black liquor.Customarily, IRD is expressed as the percentage, on a massbasis, of the as-fired BLDS (Eq. [1]):

(1)

where M As-fired and M Virgin are the mass flow rates of as-firedand virgin BLDS, respectively. Most kraft pulp mills monitorthe volume flow rates and solids content of the as-fired blackliquor online and use these values to calculate the mass flowrate, M As-fired , assuming a black liquor density value. Althoughthe liquor density changes with solids content [8], some millsuse only one fixed value for a wide range of their solids con-tents. This inherently results in an error in the M As-fired calcula-tion, particularly during the time when there is a wide varia-tion in liquor solids content and IRD load. A more proper wayto calculate as-fired black liquor mass flow rates must take intoaccount the change in liquor density with solids content.Furthermore, most mills do not continuously measure the

volume flow rate and/or solids content of virgin black liquor,so M Virgin is usually not available. As a result, although Eq. (1)is simple, it cannot be used readily.

Dust dissolution methodSeveral mills measure the amount of IRD on a daily basis bydirecting the precipitator dust screw feeder to a small, dedi-cated tank filled with a fixed amount of water for a fixed pe-riod and analyzing the concentration and composition of thedissolved dust in the tank solution. The analysis results arethen used to back-calculate the amount of IRD. The principleof this method is similar to that of the online recovery boilerdust analyzer used by Tamminen et al. [6,7] mentioned above,except that it operates on a much larger scale. While the meth-od is relatively accurate as it deals directly with the duststream, the procedure is complicated and the equipment in-

volved is rather costly and requires frequent maintenance. Asa result, it has not been widely adopted.

In mills equipped with an ash treatment system to prefer-entially remove chloride (Cl) and potassium (K) from recoveryboiler precipitator ash [9], the amount of ash may be knownthrough the material balance around the ash treatment sys-tem. However, this method is limited only to mills where theash comes from one boiler and all is treated.

THEORETICAL CONSIDERATION

The amount of IRD can be determined by performing a mate-rial balance around the recovery boiler. Consider a systeminvolving three streams, in which stream 1 mixes with stream2 to form stream 3, as shown in Fig. 2 . M1, M2, and M 3 are themass flows of streams 1, 2, and 3, respectively (Eq. [2]):

2. System with three streams.


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M1 + M2 = M3 (2)

In Eqs. (3) and (4), if “a” and “b” are two particular species instream 1, stream 2, and stream 3, then:

M1Ca1 + M2Ca2 = M3Ca3 (3)

and

M1Cb1 + M2Cb2 = M3Cb3 (4)

where C a1, Ca2, and C a3 are concentrations, in wt%, of “a” instreams 1, 2, and 3, respectively, and C b1, Cb2, and C b3 are con-centrations, in wt%, of “b” in streams 1, 2, and 3, respectively.Rearranging the equations yields Eq. (5):

(5)

We now consider a case around a recovery boiler as shown inFig. 3 , where M Virgin BL, MDust , M As-fi red BL, MSmelt , and M Stack ,respectively, represent the mass flows of virgin black liquor,precipitator dust, as-fired black liquor, smelt, and stack gas.

Applying Eqs. (1)–(5) to the black l iquor mix tank, theamount of IRD, expressed as % as-fired black liquor, can becalculated (Eq. [6]):

(6)

where C a As-fired BL, Ca Virgin BL , and C a Dust , respectively, are con-centrations in wt% of species “a” in as-fired black liquor, virginblack liquor, and precipitator dust. Equation (6) suggests thereis no need to have the mass flow rates of various streams forthe calcul

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