JP3632272B2 - Rotor shaft for steam turbine and its manufacturing method, steam turbine power plant and its steam turbine - Google Patents

Description

本発明の１２％Ｃｒ系耐熱鋼においては、特に６２１℃以上の蒸気中で使用される場合には、６２５℃，１０^５ｈクリープ破断強度１０ｋｇｆ／ｍｍ^２以上，室温衝撃吸収エネルギー１ｋｇｆ−ｍ以上にすることが好ましい。
【００５６】
（１）本発明における低圧蒸気タービンの最終段ブレードに用いる１２％Ｃｒ系鋼の好ましい成分範囲限定理由について説明する。
【００５７】
Ｃは高い引張強さを得るために最低０．０８ ％必要である。あまりＣを多くすると、靭性を低下させるので０．２０ ％以下にしなければならない。特に、０．１０〜０．１８ ％が好ましい。より０．１２〜０．１６％が好ましい。
【００５８】
Ｓｉは脱酸剤、Ｍｎは脱硫酸・脱酸剤で鋼の溶解の際に添加するものであり、少量でも効果がある。Ｓｉはδフェライト生成元素であり、多量の添加は、疲労及び靭性を低下させる有害なδフェライト生成の原因になるので、０．２５％ 以下にしなければならない。なお、カーボン真空脱酸法及びエレクトロスラグ溶解法などによればＳｉ添加の必要がなく、Ｓｉ無添加がよい。特に、０．１０ ％以下、より０．０５％以下が好ましい。
【００５９】
多量のＭｎは靭性を低下させるので、０．９％ 以下にすべきである。特に、 Ｍｎは脱酸剤として有効なので、靭性向上の点から０．４％以下、より０．２％以下が好ましい。
【００６０】
Ｃｒは耐食性と引張強さを高めるが、１３％以上添加するとδフェライト組織生成の原因になる。８％より少ないと耐食性と引張強さが不十分なので、Ｃｒは８〜１３％に決定された。特に強度の点から１０．５〜１２．５％が、より１１〜１２％好ましい。
【００６１】
Ｍｏは固溶強化及び析出強化作用によって引張強さを高める効果がある。Ｍｏは引張強さ向上効果が不十分であり３％以上になるとδフェライト生成原因になるので１．５〜３．０％に限定される。特に、１．８〜２．７％、より２．０〜２．５％好ましい。なお、Ｗ及びＣｏもＭｏと同じ様な効果がある。
【００６２】
Ｖ及びＮｂは炭化物を析出し引張強さを高めると同時に靭性向上効果がある。Ｖ０．０５％，Ｎｂ０．０２％以下ではその効果が不十分であり、Ｖ０．３５％，Ｎｂ０．２％以上ではδフェライト生成の原因となる。特にＶは０．１５〜０．３０％、より０．２５〜０．３０％、Ｎｂは０．０４〜０．１５％、より０．０６〜０．１２％が好ましい。Ｎｂの代わりにＴａを全く同様に添加でき、複合添加することができる。
【００６３】
Ｎｉは低温靭性を高めと共に、δフェライト生成の防止効果がある。この効果は、Ｎｉ２％以下では不十分で、３％を越えると添加で効果が飽和する。特に
、２．３〜２．９％が好ましい。より好ましくは２．４〜２．８％である。
【００６４】
Ｎは引張強さの向上及びδフェライトの生成防止に効果があるが０．０２％ 未満ではその効果が十分でなく、０．１％ を越えると靭性を低下させる。特に、 ０．０４〜０．０８、より０．０６〜０．０８％の範囲で優れた特性が得られる。
【００６５】
Ｓｉ，Ｐ及びＳの低減は、引張強さを損なわず、低温靭性を高める効果があり、極力低減することが望ましい。低温靭性向上の点からＳｉ０．１％以下，Ｐ ０．０１５％以下，Ｓ０．０１５％以下が好ましい。特に、Ｓｉ０．０５％以下，Ｐ０．０１０％以下，Ｓ０．０１０％以下が望ましい。Ｓｂ，Ｓｎ及びＡｓの低減も、低温靭性を高める効果があり、極力低減することが望ましいが、現状製鋼技術レベルの点から、Ｓｂ０．００１５％以下，Ｓｎ０．０１％以下、及びＡｓ０．０２％以下に限定した。特に、Ｓｂ０．００１％，Ｓｎ０．００５ ％及びＡｓ０．０１％以下が望ましい。
【００６６】
さらに、本発明においては、Ｍｎ／Ｎｉ比を０．１１ 以下にするのが好ましい。
【００６７】
本発明材の熱処理は、まず完全なオーステナイトに変態するに十分な温度，最低１０００℃，最高１１００℃に均一加熱し、急冷し（好ましくは油冷）、次いで５５０〜５７０℃の温度に加熱保持・冷却し（第１次焼戻し）、次いで５６０〜６８０℃の温度に加熱保持し第２次焼戻しを行い、全焼戻しマルテンサイト組織とするものが好ましい。
【００６８】
（２）本発明における６２０〜６４０℃蒸気タービンの高圧と中圧又は高中圧一体型のロータシャフト胴部，ブレード，ノズル，内部ケーシング締付ボルト及び中圧部初段ダイヤフラムを構成するフェライト系耐熱鋼の組成の好ましい限定理由について説明する。
【００６９】
Ｃは焼入れ性を確保し、焼戻し熱処理過程で炭化物を析出させて高温強度を高めるのに不可欠の元素であり、また高い引張強さを得るためにも０．０５ ％以上必要な元素であるが、０．２０ ％を越えると高温に長時間さらされた場合に金属組織が不安定になり長時間クリープ破断強度を低下させるので、０．０５ 〜０．２０％に限定される。望ましくは０．０６〜０．１４％、又は０．０８〜０．１３％であり、特に０．０９〜０．１２％が好ましい。
【００７０】
Ｍｎは脱酸剤等のために添加するものであり、少量の添加でその効果は達成され、１．５％ を越える多量の添加はクリープ破断強度を低下させるので好ましくない。特に０．０３〜０．２０％又は０．３〜０．７％が好ましく、多い方に対しては０．３５〜０．６５％がより好ましい。Ｍｎの少ない方が高強度が得られる。また、Ｍｎ量の多い方は加工性がよい。
【００７１】
Ｓｉも脱酸剤として添加するものであるが、真空Ｃ脱酸法などの製鋼技術によれば、Ｓｉ脱酸は不要である。Ｓｉを低くすることにより有害なδフェライト組織生成防止と結晶粒界偏析等による靭性低下を防止する効果がある。したがって、添加する場合には0.15 ％以下に抑える必要があり、望ましくは０.０７％以下であり、特に０.０４％未満が好ましい。
【００７２】
Ｎｉは靭性を高め、かつ、δフェライトの生成を防止するのに非常に有効な元素であるが、０．０５％未満ではその効果が十分でなく、１．０％を越える添加はクリープ破断強度を低下させるので好ましくない。特に０．３〜０．７％、より ０．４〜０．６５％が好ましい。
【００７３】
Ｃｒは高温強度及び高温耐酸化を高めるのに不可欠の元素であり、最低９％必要であるが、１３％を越えると有害なδフェライト組織を生成し高温強度及び靭性を低下させるので、９〜１３％に限定される。特に１０〜１２％、より１０.８〜１１.８％が好ましい。
【００７４】
Ｍｏ添加は、高温強度向上のために行われる。しかし、本発明鋼の様に１％を越えるＷを含む場合には、０．５ ％以上のＭｏ添加は靭性及び疲労強度を低下させるので、０．５％以下に制限される。特に０．０５〜０．４５％、より０．１〜 ０．２ ％が好ましい。
【００７５】
Ｗは高温での炭化物の凝集粗大化を抑制し、またマトリックスを固溶強化するので、６２０℃以上の高温長時間強度を顕著に高める効果がある。６２０℃では１〜１．５ ％、６３０℃では１．６〜２．０％、６４０℃では２．１〜２．５％、 ６５０℃では２．６〜３．０％、６６０℃では３．１〜３．５％とするのが好ましい。またＷが３．５ ％を越えるとδフェライトを生成して靭性が低くなるので、１〜３．５ ％に限定される。特に２．１〜３．０％が好ましく、より２．５〜２．８％が好ましい。
【００７６】
Ｖは、Ｖの炭窒化物を析出してクリープ破断強度を高める効果があるが、０．０５％未満ではその効果が不十分で０．３ ％を越えるとδフェライトを生成して疲労強度を低下させる。特に０．１０〜０．２５％が好ましく、より０．１５〜０．２３％が好ましい。
【００７７】
ＮｂはＮｂＣ炭化物を析出し、高温強度を高めるのに非常に効果的な元素であるが、あまり多量に添加すると、特に大型鋼塊では粗大な共晶ＮｂＣ炭化物が生じ、かえって強度を低下させたり、疲労強度を低下させるδフェライトを析出させる原因になるので０．２０％以下に抑える必要がある。また０．０１％未満の Ｎｂでは効果が不十分である。特に０．０２〜０．１５％、又は０．０３〜０．１０％、より０．０４〜０．０８％が好ましい。
【００７８】
Ｃｏは本発明を従来の発明から区別して特徴づける重要な元素である。本発明においては、Ｃｏ添加により高温強度が著しく改善されるとともに、靭性も高める。これは、Ｗとの相互作用によると考えられ、Ｗを１％以上含む本発明合金において特徴的な現象である。このようなＣｏの効果を実現するために、本発明合金におけるＣｏの下限は２．０ ％であるが、過度に添加してもより大きな効果が得られないだけでなく、延性が低下するので、上限は１０％になる。望ましくは６２０℃に対しては２〜３％、６３０℃に対しては３．５〜４．５％、６４０℃に対しては５〜６％、６５０℃に対しては６．５〜７．５％、６６０℃に対しては８〜９％が望ましい。
【００７９】
Ｎも本発明を従来の発明から区別して特徴づける重要な元素である。Ｎはクリープ破断強度の改善及びδフェライト組織の生成防止に効果があるが０．００５ ％以下ではその効果が十分でなく０．０５ ％を越えると靭性を低下させると共に、クリープ破断強度も低下させる。特に０．０１〜０．０３％が、より０．０１５ 〜０．０２５％が好ましい。
【００８０】
Ｂは粒界強度作用とＭ_２３Ｃ_６炭化物中に固溶し、Ｍ_２３Ｃ_６型炭化物の凝集粗大化を妨げる作用により高温強度を高める効果があり、０．００１ ％を越える添加が有効であるが、０．０３％を越えると溶接性や鍛造性を害するので、０．００１〜０．０３ ％に制限される。望ましくは０．００１〜０．０１％、又は０．０１〜 ０．０２％が好ましい。
【００８１】
Ｔａ，Ｔｉ及びＺｒの添加は、靭性を高める効果があり、Ｔａ０．１５％以下，Ｔｉ０．１％以下及びＺｒ０．１％以下の単独または複合添加で十分な効果が得られる。Ｔａを０．１ ％以上添加した場合にはＮｂの添加を省略することができる。
【００８２】
本発明におけるロータシャフト胴部及び動翼と静翼の少なくとも初段は６２０〜６３０℃の蒸気温度に対してはＣ０．０９〜０．２０％，Ｓｉ０．１５ ％以下，Ｍｎ０．０５〜１．０％，Ｃｒ９．５〜１２．５％，Ｎｉ０．１〜１．０％，Ｖ０．０５〜０．３０％，Ｎ０．０１〜０．０６％，Ｍｏ０．０５〜０．５％，Ｗ２〜３．５％，Ｃｏ２〜４．５％，Ｂ０．００１〜０．０３０％，７７％ 以上のＦｅを有する全焼戻しマルテンサイト組織を有する鋼によって構成されるものが好ましい。また、６３５〜６６０℃の蒸気温度に対しては前述のＣｏ量を５〜８％とし、７８％以上のＦｅを有する全焼戻しマルテンサイト組織を有する鋼によって構成されるのが好ましい。特に、両者の温度に対してＭｎ量を０．０３〜０．２％及びＢ量を ０．００１〜０．０１％と少なくすることによって高強度が得られる。特に、Ｃ ０．０９〜０．２０％，Ｍｎ０．１〜０．７％，Ｎｉ０．１〜１．０％，Ｖ０．１０〜 ０．３０％，Ｎ０．０２〜０．０５％，Ｍｏ０．０５〜０．５％，Ｗ２〜３．５％を含有し、６３０℃以下に対してはＣｏ２〜４％，Ｂ０．００１〜０．０１％及び６３０〜６６０℃に対してはＣｏ５．５〜９．０％，Ｂ０．０１〜０．０３％とするのが好ましい。
【００８３】
前述の式によって求められるＣｒ当量をロータシャフトに対しては４〜１０．５、特に６．５〜９．５が好ましく、他のものも同様である。
【００８４】
本発明の蒸気タービンの高圧と中圧のロータ材は、δフェライト組織が混在すると、疲労強度及び靭性が低くなるので、組織は均一な焼戻しマルテンサイト組織が好ましい。焼戻しマルテンサイト組織を得るために、前述の式で計算されるＣｒ当量を、成分調整により１０以下にしなければならない。Ｃｒ当量をあまり低くするとクリープ破断強度が低下してしまうので、４以上が好ましく、特に、Ｃｒ当量５〜８が好ましい。
【００８５】
ロータ胴部は、６２１℃以上の蒸気中で高速回転（３０００又は３６００ｒｐｍ）されるので、ブレードを支持しているダブテール部と中心孔部には、高い応力が作用するので、クリープ破壊防止の観点から、１０ｋｇｆ／ｍｍ^２以上の１０^５ｈクリープ破壊強度が要求される。また、起動時にはメタル温度が低い時に中心孔部に引張り熱応力が作用するので、脆性破壊防止の観点から、１ｋｇｆ−ｍ以上の室温衝撃吸収エネルギーが要求される。
【００８６】
（３）次にロータジャーナル部及び低温域部の好ましい成分限定理由とその働きについて説明する。
【００８７】
Ｃは高い引張強さを得るために０．０６％以上必要な元素であるが、０．１４％を越えると溶接性を悪くするので、０．０６〜０．１４％が好ましい。特に０．０８〜０．１１％が好ましい。
【００８８】
Ｎはδフェライト組織の生成防止に効果があるが、０.００５％ 以下ではその効果が十分でなく０ . ０５％を越えると靭性を低下させるとともに溶接性を悪くする。特に、0.015〜０.０３％が好ましい。
【００８９】
Ｍｎは脱硫酸として添加するものであり、２％以下の添加でその効果は達成される。特に１％以下が好ましい。
【００９０】
Ｓｉも脱酸剤として添加するものであるが、真空Ｃ脱酸法などの製鋼技術によれば、Ｓｉ脱酸は不要である。またＳｉを低くすることにより有害なδフェライト組織生成防止効果がある。したがって、添加する場合には０．５ ％以下に抑えることが好ましく、特に０．２ ％以下が好ましい。
【００９１】
Ｖは焼入れ性を高める効果があるが、０．０５ ％以下ではその効果が不十分で０．３５ ％を越えるとδフェライトを生成して疲労強度を低下させる恐れがある。特に、０．１５〜０．２５％が好ましい。
【００９２】
Ｎｂは靭性を高めるのに効果的な元素であるが、あまり多量に添加すると、特に大型鋼塊では粗大な共晶Ｎｂ炭化物が生じ、かえって靭性を低下させたり、疲労強度を低下させるδフェライトを析出させる恐れがあるので０．２ ％以下に抑えることが好ましい。また０．０１ ％以下のＮｂでは効果が不十分である。特に大型鋼塊の場合は０．０３〜０．１％が、より０．０４〜０．０８％が好ましい。
【００９３】
Ｎｉは靭性を高め、かつ、δフェライト生成を防止するのに非常に有効な元素であるが、０．２ ％以下ではその効果が十分でなく、２％を越える添加は残留オーステナイト組織を生成させるので好ましくない。
【００９４】
Ｃｒは高強度及び高温酸化を改善する効果がある。１３％を越えると有害なδフェライト組織生成の原因となり、７％より少ないと高温高圧蒸気に対する耐酸化性が不十分となる。また過剰のＣｒ添加は有害なδフェライト組織生成及び靭性低下の原因となる。特に、９．５ 〜１２％、より１０．５〜１１．５％が好ましい。
【００９５】
Ｗは焼戻し抵抗高め引張強さを高める効果がある。しかし、Ｗが３％を越えると靭性が低くなる。２．０〜２．８％が好ましい。
【００９６】
Ｍｏも焼戻し抵抗高め引張強さを高める効果がある。１％以上のＭｏ添加は靭性及び疲労強度を低下させる恐れがあるので、１％以下が好ましい。
【００９７】
Ｔａ，Ｔｉ及びＺｒの添加は、靭性を高める効果があり、Ｔａ０．２ ％以下，Ｔｉ０．１％以下及びＺｒ０．１％以下の単独または複合添加で十分な効果が得られる。Ｔａを添加した場合には、Ｎｂの添加を省略することができ、特に０．１％以上が好ましい。
【００９８】
Ｂ添加は引張強さを高める効果がある。Ｂ含有量が０．００３０％を越えると、溶接性が著しく悪くなる恐れがあり、上限は０．００３０％、より０．００２０％が好ましい。特に無添加が好ましい。
【００９９】
Ｃｏ添加は有害組織であるδフェライト相の析出を防止すると共に、靭性を高める効果がある。５％以上を越える添加は、靭性を低める恐れがある。特に１〜３％が好ましい。
【０１００】
ロータ全体を均質にするのには、鋼塊重量８０トン前後（ロータ直径：１２００ｍｍ，長さ：約８ｍ）と大型になるので、高度な製造技術が要求される。本発明ロータは目標組成とする合金原料を電気炉で溶解し精錬後、真空カーボン脱酸法で脱酸し、電極を作製し、この電極を用いエレクトロスラグ再溶解法で鋼塊を作製し、この鋼塊を熱間鍛錬で成形することにより作製できる。電気炉とエレクトロスラグ再溶解法で、２回溶解を繰り返すことにより、成分偏析の少ない均質なロータが作製できる。本発明の１２％Ｃｒ系耐熱鋼製ロータシャフトは、溶接性良好な合金鋼及び高温強度の高い合金鋼の２種以上の消耗電極を上述のように別々に準備し、まずジャーナル部に該当する前者の消耗電極をエレクトロスラグ溶解した後、直ちに胴部に該当する後者の消耗電極をエレクトロスラグ溶解して接合し、その後再び前者の消耗電極をエレクトロスラグ溶解し継ぎ足すことによって一体の所定の長さと直径の鋼塊が得られる。また、ジャーナルを溶接性良好な合金鋼（上端部及び下端部）で胴部（中央部）を高温強度の高い合金鋼で作製した消耗電極を準備し、この消耗電極をエレクトロスラグ溶解することによっても得られる。
【０１０１】
また、前記鋼塊の熱間鍛錬は、高温で行うほど変形抵抗が小さく鍛錬し易いが、あまり高温で行うと割れてしまうので、８５０〜１１５０℃の温度範囲で行わなければならない。熱間鍛錬成形後、９５０〜１１５０℃に加熱し焼鈍し、更に１０００〜１１５０℃に加熱焼入れ後、５５０〜６５０℃及び６５０〜７５０℃で２回焼戻しを行うことにより、１０ｋｇｆ／ｍｍ^２以上の６５０℃，１０^５ｈクリープ破断強度と１ｋｇｆ−ｍ以上の室温衝撃吸収エネルギーが得られ、６２１℃以上の蒸気中で使用可能な蒸気タービンロータが製造法できる。また、２回焼戻しは、残留オーステナイトを完全に分解させ、均一な全焼戻しマルテンサイト組織にすることができる。
【０１０２】
（４）本発明に係るロータシャフトはジャーナル部表面に軸受特性の高い低合金鋼を肉盛溶接によって所望の厚さの肉盛溶接層を形成させるものである。
【０１０３】
本発明における１２％Ｃｒ系マルテンサイト鋼からなる蒸気タービンロータシャフトのジャーナル部に形成される軸受特性の高い肉盛溶接層は好ましくは鋼からなる溶接材を用いて５層〜１０層の前記肉盛溶接層を形成し、初層から２層目〜４層目のいずれかまでの前記溶接材のＣｒ量を順次低下させるとともに、４層目以降を同じＣｒ量を有する鋼からなる溶接材を用いて溶接し、前記初層の溶接に用いられる溶接材のＣｒ量を前記母材のＣｒ量より２〜６％程度少なくし、４層目以降の溶接層のＣｒ量を０．５〜３％（好ましくは１〜２．５％）とするものである。
【０１０４】
本発明においては、ジャーナル部の軸受特性の改善には肉盛溶接が最も安全性が高い点で好ましいものであるが、その肉盛溶接は鋼中のＢ量の増加によってきわめて困難になるが、ジャーナル部のＢ量を低くすることにより胴部をより高強度とするためにＢ量を０．００５〜０．０２％含有させることができる。
【０１０５】
本発明法によって得られる肉盛溶接層は５層〜１０層とするのが好ましい。前述の如く、初層溶接層としてＣｒ量の急激な低下は高い引張残留応力の発生、或いは溶接割れ発生の原因となることからその溶接材としてのＣｒ量を大幅に減らすことができないので、溶接層数を多くして徐々にＣｒ量を下げるが、更に表面層として所望のＣｒ量をその所望の厚さとを確保することから５層以上とする。尚、１０層以上溶接してもそれ以上の効果は得られない。蒸気タービンロータシャフトの如く大型構造材としては、肉盛溶接層として母材からの組成の影響を受けず、かつ所望の組成と所望の厚さとし、母材の影響のない厚さとして３層及びその上に所望の特性のものを所望の厚さとし、その厚さとして２層以上、一例として最終仕上げで約１８ｍｍの厚さとする。このような厚さを形成するには切削による最終仕上げ代を除いても５層の肉盛溶接層が好ましい。３層目以降は主に焼戻しマルテンサイト組織を有し、炭化物が析出していることが好ましい。特に、４層目以降の溶接層の組成として重量で、Ｃ０．０１〜０．１％，Ｓｉ０．３ 〜１％，Ｍｎ０．３〜１．５％，Ｃｒ０．５〜３％，Ｍｏ０．１〜１．５ ％を含み残部 Ｆｅからなるものが好ましい。
【０１０６】
また、肉盛溶接層は初層より２層目〜４層目のいずれかまでを順次Ｃｒ量を低下させるもので、肉盛溶接にあたって層毎に徐々にＣｒ含有量を低めた溶接棒を用いて溶接すれば、初層溶接部のクロム含有量の大幅な違いによる初層溶接部の延性低下の問題が生ぜず、溶接割れを生じることなく所望の組成の肉盛溶接層を形成することができる。これにより、本発明は母材と初層部付近のクロム含有量が極端に差を示すことなく、しかも最終層に上述の軸受特性の高い肉盛溶接層を形成することができる。
【０１０７】
初層溶接に適用する溶接材としてはそのクロム含有量を母材のクロム量より２〜６重量％程度少なくする。溶接材のＣｒ量を母材より低い値として２％以下では肉盛溶接層のＣｒ量を十分に下げることができず、効果が小さい。逆に、６％以上では母材と肉盛溶接層との急激なＣｒ量の低下につながり、このＣｒ量の差が熱膨張係数の差を生じ高い引張残留応力の発生、或いは溶接割れ発生の原因となる。尚、高Ｃｒほど熱膨張係数が小さいので、低Ｃｒとなる肉盛溶接層は母材より熱膨張係数が大きく溶接後に高い引張残留応力が形成される。そのためより低Ｃｒ鋼での溶接は高い残留応力のため硬さが高く、また溶接割れ発生の原因となるので、溶接材のＣｒ量は母材のそれより少ない値として６％以下とする必要がある。このような溶接材を使用することにより初層溶接部のクロム含有量は母材と混合するため、母材よりも約１〜３％低くなる程度にとどまり、良好な溶接が得られる。
【０１０８】
本発明法において、４層以降を同じＣｒ量を有する鋼からなる溶接材を用いて形成する。肉盛溶接において、３層目までは母材の組成の影響を受けるが、４層目以降の肉盛溶接層の組成は用いられる溶接材の組成によってのみ形成されるので、蒸気タービンロータシャフトのジャーナル部として必要な特性を満たすものを形成させることができる。従って、前述のように蒸気タービンロータシャフトとしての大型構造物として必要な肉盛溶接層は約１８ｍｍであるので、最終層として必要な合金組成とその組成での必要な十分な厚さを確保するために４層目以降を同じＣｒ量の溶接材によって２層以上溶接することになり前述のジャーナル部として要求される特性を満足するものを十分な厚さをもって形成させることができる。
【０１０９】
本発明におけるブレード，ノズル，内部ケーシング締付ボルト，中圧部初段ダイヤフラムは真空溶解によって溶解され、真空下で金型に鋳造され、インゴットが製造される。インゴットは前述と同様の温度で所定形状に熱間鍛造され、１０５０〜１１５０℃で加熱後水冷又は油焼入れされ、次いで７００〜８００℃で焼戻し処理が施され、切削加工によって所望の形状のブレードとなる。真空溶解は１０^−１〜１０^−４ｍｍＨｇ下で行われる。特に、本発明における耐熱鋼は高圧部及び中圧部のブレード及びノズルの全段に用いることができるが、特に、両者の初段には必要なものである。
【０１１０】
３層目以降は主に焼戻しマルテンサイト組織を有し、炭化物が析出していることが好ましい。特に、４層目以降の溶接層の組成として重量で、Ｃ０．０１〜０．１％，Ｓｉ０．３〜１％，Ｍｎ０．３〜１．５％，Ｃｒ０．５〜３％，Ｍｏ０．１〜１．５％を含み残部Ｆｅからなるものが好ましい。
【０１１１】
（５）本発明の高圧タービン，中圧タービン及び高中圧タービンの内部ケーシング加減弁弁箱，組合せ再熱弁弁箱，主蒸気リード管，主蒸気入口管，再熱入口管，高圧タービンノズルボックス，中圧タービン初段ダイヤフラム，高圧タービン主蒸気入口フランジ，エルボ，主蒸気止め弁を構成するフェライト系耐熱鋼の組成の限定理由について説明する。
【０１１２】
フェライト系耐熱鋳鋼ケーシング材においては、特にＮｉ／Ｗ比を０．２５〜０．７５に調整することにより、６２１℃，２５０ｋｇｆ／ｃｍ^２以上の超々臨界圧タービン高圧及び中圧内部ケーシング並びに主蒸気止め弁及び加減弁ケーシングに要求される、６２５℃，１０^５ｈクリープ破断強度９ｋｇｆ／ｍｍ^２以上，室温衝撃吸収エネルギー１ｋｇｆ−ｍ以上の耐熱鋳鋼ケーシング材が得られる。
【０１１３】
本発明のフェライト系耐熱鋳鋼ケーシング材においては、高い高温強度と低温靭性並びに高い疲労強度を得るために、前述の式で計算されるＣｒ当量を４〜 １０に成分調整することが好ましい。
【０１１４】
本発明の１２％Ｃｒ系耐熱鋼においては、６２１℃以上の蒸気中で使用されるので、６２５℃，１０^５ｈクリープ破断強度９ｋｇｆ／ｍｍ^２以上，室温衝撃吸収エネルギー１ｋｇｆ−ｍ以上にしなければならない。更に、より高い信頼性を確保するためには、６２５℃，１０^５ｈクリープ破断強度１０ｋｇｆ／ｍｍ^２以上，室温衝撃吸収エネルギー２ｋｇｆ−ｍ以上であることが好ましい。
【０１１５】
Ｃは高い引張強さを得るために０．０６％以上必要な元素であるが、０．１６％を越えると高温に長時間さらされた場合に金属組織が不安定になり長時間クリープ破断強度を低下させるので、０．０６〜０．１６％に限定される。特に０．０９ 〜０．１４ ％が好ましい。
【０１１６】
Ｎはクリープ破断強度の改善及びδフェライト組織の生成防止に効果があるが、０．０１％未満ではその効果が十分でなく、０．１％を越えても顕著な効果はなく、逆に靭性を低下させると共に、クリープ破断強度も低下させる。特に０．０２〜０．０６ ％が好ましい。
【０１１７】
Ｍｎは脱酸剤として添加するものであり、少量の添加でその効果は達成され、１％を越える多量の添加はクリープ破断強度を低下させ、特に０．４〜０．７％が好ましい。
【０１１８】
Ｓｉも脱酸剤として添加するものであるが、真空Ｃ脱酸法などの製鋼技術によれば、Ｓｉ脱酸は不要である。またＳｉを低くすることにより有害なδフェライト組織生成防止効果がある。したがって、添加する場合には０．５ ％以下に抑える必要があり、特に０．１〜０．４％が好ましい。
【０１１９】
Ｖはクリープ破断強度を高める効果があるが、０．０５ ％未満ではその効果が不十分で０．３５ ％を越えるとδフェライトを生成して疲労強度を低下させる。特に、０．１５〜０．２５％が好ましい。
【０１２０】
Ｎｂは高温強度を高めるのに非常に効果的な元素であるが、あまり多量に添加すると、特に大型鋼塊では粗大な共晶Ｎｂ炭化物が生じ、かえって強度を低下させたり、疲労強度を低下させるδフェライトを析出させる原因になるので０．１５％以下に抑える必要がある。また０．０１ ％未満のＮｂでは効果が不十分である。特に大型鋼塊の場合は０．０２〜０．１％が、より０．０４〜０．０８が好ましい。Ｎｉは靭性を高め、かつ、δフェライトの生成を防止するのに非常に有効な元素であるが、０．２％未満ではその効果が十分でなく、１．０％を越える添加はクリープ破断強度を低下させるので好ましくない。特に０．４〜０．８％が好ましい。
【０１２１】
Ｃｒは高強度及び高温酸化を改善する効果がある。１２％を越えると有害なδフェライト組織生成の原因となり、８％より少ないと高温高圧蒸気に対する耐酸化性が不十分となる。またＣｒ添加は、クリープ破断強度を高める効果があるが、過剰の添加は有害なδフェライト組織生成及び靭性低下の原因となる。特に ８．０ 〜１０％、より８．５〜９．５％が好ましい。
【０１２２】
Ｗは高温長時間強度を顕著に高める効果がある。１％より少ないＷでは、６２０〜６６０℃で使用する耐熱鋼としては効果が不十分である。またＷが４％を越えると靭性が低くなる。６２０℃では１．０〜１．５％、６３０℃では１．６〜２．０％、６４０℃では２．１〜２．５％、６５０℃に対しては２．６〜３．０％、６６０℃では３．１〜３．５％が好ましい。
【０１２３】
ＷとＮｉとは互いに相関性があり、Ｎｉ／Ｗ比を０．２５〜０．７５とすることにより強度と靭性ともに高いものが得られる。
【０１２４】
Ｍｏ添加は、高温強度向上のために行われる。しかし、本発明鋳鋼の様に１％を越えるＷを含む場合には、１．５ ％以上のＭｏ添加は靭性及び疲労強度を低下させるので、１．５ ％以下がよく、特に０．４〜０．８％、より０．５５〜０．７０％が好ましい。
【０１２５】
Ｔａ，Ｔｉ及びＺｒの添加は、靭性を高める効果があり、Ｔａ０．１５％以下，Ｔｉ０．１％以下及びＺｒ０．１％以下の単独または複合添加で十分な効果が得られる。Ｔａを０．１ ％以上添加した場合には、Ｎｂの添加を省略することができる。
【０１２６】
本発明の耐熱鋳鋼ケーシング材は、δフェライト組織が混在すると、疲労強度及び靭性が低くなるので、組織は均一な焼戻しマルテンサイト組織が好ましい。焼戻しマルテンサイト組織を得るために、前述の式で計算されるＣｒ当量を、成分調整により１０以下にしなければならない。Ｃｒ当量をあまり低くするとクリープ破断強度が低下してしまうので、４以上にしなければならない。特に、Ｃｒ当量６〜９が好ましい。
【０１２７】
Ｂ添加は高温（６２０℃以上）クリープ破断強度を著しく高める。Ｂ含有量が０．００３％を越えると、溶接性が悪くなるため、上限は０．００３％に制限される。特に、大型ケーシングのＢ含有量の上限は０．００２８％、更に０．０００５〜０．００２５ ％が好ましく、特に０．００１〜０．００２％が好ましい。
【０１２８】
ケーシングは、６２０℃以上の高圧蒸気をカバーしているので、内圧による高応力が作用する。そのため、クリープ破壊防止の観点から、１０ｋｇｆ／ｍｍ^２ 以上の１０^５ ｈクリープ破断強度が要求される。また、起動時には、メタル温度が低い時に熱応力が作用するので、脆性破壊防止の観点から、１ｋｇｆ−ｍ以上の室温衝撃吸収エネルギーが要求される。より高温度側に対してはＣｏを１０％以下含有させることにより強化が図れる。特に、６２０に対しては１〜２％、６３０℃に対しては２．５〜３．５％，６４０℃に対しては４〜５％、６５０℃に対しては５．５〜６．５％、６６０℃に対しては７〜８％が好ましい。６００〜６２０℃では無添加でもよい。
【０１２９】
欠陥の少ないケーシングを作製するには、鋳塊重量５０トン前後と大型になるので、高度な製造技術が要求される。本発明フェライト系耐熱鋳鋼ケーシング材は、目標組成とする合金原料を電気炉で溶解し、とりべ精錬後、砂型鋳型に鋳込み成形することにより健全なものが作製できる。鋳込み前に、十分な精錬及び脱酸を行うことにより、引け巣等の鋳造欠陥の少ないものにできる。
【０１３０】
また、前記の鋳鋼を１０００〜１１５０℃で焼鈍熱処理後、１０００〜１１００℃に加熱し急冷する焼準熱処理，５５０〜７５０℃及び６７０〜７７０℃の順序で２回焼戻しを行うことにより、６２１℃以上の蒸気中で使用可能な蒸気タービンケーシングが製造できる。焼鈍及び焼準温度は、１０００℃以下では炭窒化物を十分固溶させることができず、あまり高くすると結晶粒粗大化の原因になる。また、２回焼戻しは、残留オーステナイトを完全に分解させ、均一な焼戻しマルテンサイト組織にすることができる。上記の製法で作製することにより、１０ｋｇｆ／ｍｍ^２ 以上の６２５℃，１０^５ ｈクリープ破断強度と１ｋｇｆ−ｍ以上の室温衝撃吸収エネルギーが得られ、６２０℃以上の蒸気中で使用可能な蒸気タービンケーシングにできる。
【０１３１】
Ｏは０．０１５％を越えると高温強度及び靭性値を低下させるので、０．０１５％以下が好ましく、特に０．０１０ ％以下が好ましい。
【０１３２】
本発明におけるケーシングは前述のＣｒ当量とし、δフェライト量が５％以下にするのが好ましく、より０％がよい。
【０１３３】
内部ケーシングを鋳鋼によって製造する他は鍛鋼によって製造するのが好ましい。
【０１３４】
（６）低圧蒸気タービンロータシャフトは重量で、Ｃ０．２〜０．３％，Ｓｉ０．１％以下，Ｍｎ０．２％以下，Ｎｉ３．２〜４．０％，Ｃｒ１．２５〜２．２５％，Ｍｏ０．１〜０．６％，Ｖ０．０５〜０．２５％を有する全焼戻しベーナイト組織を有する低合金鋼が好ましく、前述の高圧，中圧ロータシャフトと同様の製法によって製造されるのが好ましい。特に、Ｓｉ量は０．０５％以下，Ｍｎ０．１％以下の他Ｐ，Ｓ，Ａｓ，Ｓｂ，Ｓｎ等の不純物を極力低めた原料を用い、総量０．０２５ ％以下とするように用いられる原材料の不純物の少ないものを使用するスーパークリーン化した製造とするのが好ましい。Ｐ，Ｓ各０．０１０％以下，Ｓｎ，Ａｓ０．００５％以下，Ｓｂ０．００１％以下が好ましい。
【０１３５】
（７）低圧タービン用ブレードの最終段以外及びノズルは、Ｃ０．０５〜０．２％，Ｓｉ０．１〜０．５％，Ｍｎ０．２〜１．０％，Ｃｒ１０〜１３％，Ｍｏ０．０４〜０．２ ％を有する全焼戻しマルテンサイト鋼が好ましい。
【０１３６】
（８）低圧タービン用内部及び外部ケーシングともにＣ０．２〜０．３％，Ｓｉ ０．３〜０．７％，Ｍｎ１％以下を有する炭素鋳鋼が好ましい。
【０１３７】
（９）主蒸気止め弁ケーシング及び蒸気加減弁ケーシングはＣ０．１〜０．２％，Ｓｉ０．１ 〜０．４％，Ｍｎ０．２〜１．０％，Ｃｒ８．５〜１０．５％，Ｍｏ０．３〜１．０％，Ｗ１．０〜３．０％，Ｖ０．１〜０．３％，Ｎｂ０．０３〜０．１ ％，Ｎ０．０３〜０．０８％，Ｂ０．０００５〜０．００３％を含む全焼戻しマルテンサイト鋼が好ましい。
【０１３８】
（１０）低圧タービンの最終段動翼として１２％Ｃｒ系鋼のほかＴｉ合金が用いられ、特に４０インチを越える長さに対してはＡｌ５〜８％及びＶ３〜６％を有するＴｉ合金が用いられる。特に、４３インチにおいてはＡｌ５．５〜６．５％，Ｖ３．５〜４．５％とし、４６インチではＡｌ４〜７％，Ｖ４〜７％及びＳｎ１〜３％を有する高強度材がよい。
【０１３９】
（１１）高圧タービン，中圧タービン及び高中圧タービン用外部ケーシングにはＣ０．１０〜０．２０％，Ｓｉ０．０５〜０．６％，Ｍｎ０．１〜１．０％，Ｎｉ０．１〜０．５％，Ｃｒ１〜２．５％，Ｍｏ０．５〜１．５％，Ｖ０．１〜０．３５％を含み、好ましくはＡｌ０．０２５％以下，Ｂ０．０００５〜０．００４％及びＴｉ０．０５〜０．２ ％の少なくとも一方を含み、全焼戻しベーナイト組織を有する鋳鋼によって製造するのが好ましい。特に、Ｃ０．１０〜０．１８％，Ｓｉ０．２０〜０．６０％，Ｍｎ０．２０〜０．５０％，Ｎｉ０．１〜０．５％，Ｃｒ１．０〜１．５％，Ｍｏ０．９〜１．２％，Ｖ０．２〜０．３％，Ａｌ０．００１〜０．００５％，Ｔｉ０．０４５〜０．１０ ％及びＢ０．０００５〜０．００２０％を含む鋳鋼が好ましい。より好ましくはＴｉ／Ａｌ比が０．５〜１０である。
【０１４０】
（１２）蒸気温度６２５〜６５０℃における高圧，中圧，高中圧タービン（高圧側と中圧側）の初段ブレードとして重量で、Ｃ０．０３〜０．２０％（好ましくは０．０３〜０．１５％），Ｃｒ１２〜２０％，Ｍｏ９〜２０％（好ましくは１２〜２０％），Ｃｏ１２％以下（好ましくは５〜１２％），Ａｌ０．５〜１．５％，Ｔｉ１〜３％，Ｆｅ５％以下，Ｓｉ０．３％以下，Ｍｎ０．２％以下，Ｂ０．００３ 〜０．０１５％の他，Ｍｇ０．１％以下，希土類元素０．５％以下，Ｚｒ０．５％以下の１種以上を含むＮｉ基合金を用いることができる。以下については０％も含む。鍛造後、溶体化処理され、７００〜８７０℃で時効処理される。
【０１４１】
【発明の実施の形態】
（実施例１）
オイルショック後の燃料高騰を契機に、蒸気条件の向上による熱効率向上を図るため蒸気温度６００℃〜６４９℃微粉炭直接燃焼ボイラ及び蒸気タービンが要求される。このような、蒸気条件のボイラの一例を表１に示す。
【０１４２】
【表１】

【０１４３】
大容量化とともに微粉炭燃焼火炉が大型化し、１０５０ＭＷ級で火炉幅３１ｍ，火炉奥行き１６ｍ，１４００ＭＷ級で火炉幅３４ｍ，火炉奥行き１８ｍとなる。
【０１４４】
表２は蒸気温度６２５℃，１０５０ＭＷ蒸気タービンの主な仕様である。本実施例は、クロスコンパウンド型４流排気，低圧タービンにおける最終段翼長が ４３インチであり、ＨＰ−ＩＰにて３６００ｒｐｍ／ｍｉｎ及びＬＰ２台で１８００ ｒｐｍ／ｍｉｎの回転数を有し、高温部においては表に示す主な材料によって構成される。高圧部（ＨＰ）の蒸気温度は６２５℃，２５０ｋｇ／ｃｍ^２ の圧力であり、中圧部（ＩＰ）の蒸気温度は６２５℃に再熱器によって加熱され、１７０〜１８０ｋｇ／ｃｍ^２の圧力で運転される。低圧部（ＬＰ）は蒸気温度は４５０℃で入り、１００℃以下，７２２ｍｍＨｇの真空で復水器に送られる。
【０１４５】
【表２】

【０１４６】
図１は高圧及び中圧蒸気タービンの断面構成図である。高圧蒸気タービンは高圧内部車室１８とその外側の高圧外部車室１９内に高圧動翼１６を植設した高圧車軸（高圧ロータシャフト）２３が設けられる。前述の高温高圧の蒸気は前述のボイラによって得られ、主蒸気管を通って、主蒸気入口を構成するフランジ，エルボ２５より主蒸気入口２８を通り、ノズルボックス３８より初段複流の動翼に導かれる。初段は複流であり、片側に他８段設けられる。これらの動翼に対応して各々静翼が設けられる。動翼は鞍型ダブティル型式，ダブルティノン，初段翼長約３５ｍｍである。車軸間の長さは約５．２５ ｍ及び静翼部に対応する部分で最も小さい部分の直径は約６２０ｍｍであり、直径に対する長さの比は約８．５ である。
【０１４７】
ロータシャフトの初段と最終段の動翼植込み部分の幅はほぼ等しく、２段目，３〜５段目，６段目，７〜８段目の５段階で下流側に従って段階的に小さくなっており、２段目の植込み部の軸方向の幅は最終段のそれに対して０．６４ 倍の大きさである。
【０１４８】
ロータシャフトの静翼に対応する部分は動翼植込み部に対してロータシャフトの直径が小さくなっている。その部分の軸方向の幅は２段目動翼と３段目動翼との間の幅に対して最終段動翼とその手前の動翼との間の幅まで段階的に小さくなっており、後者の幅は前者の幅に対して０．８６ 倍と小さくなっている。２段目〜６段目までと、６段目〜９段目までとの２段階で小さくしたものである。
【０１４９】
本実施例においては後述する表５に示す材料を初段ブレード及びノズルを使用した他はいずれもＷ，Ｃｏ及びＢを含まない１２％Ｃｒ系鋼によって構成したものである。本実施例における動翼の翼部の長さは初段が３５〜５０ｍｍ、２段目から最終段になるに従って各段で長くなっており、特に蒸気タービンの出力によって２段から最終段までの長さが６５〜２１０ｍｍであり、段数は９〜１２段で、各段の翼部の長さは下流側が上流側に対して隣り合う長さで１．１０〜１．１５の割合で長くなっているとともに、下流側でその比率が徐々に大きくなっている。
【０１５０】
中圧蒸気タービンは高圧蒸気タービンより排出された蒸気を再度６２５℃に再熱器によって加熱された蒸気によって高圧蒸気タービンと共に発電機を回転させるもので、３６００回／ｍｉｎ の回転数によって回転される。中圧タービンは高圧タービンと同様に中圧内部車室２１と外部車室２２とを有し、中圧動翼１７と対抗して静翼が設けられる。動翼１７は６段で２流となり、中圧車軸（中圧ロータシャフト）の長手方向に対しほぼ対称に左右に設けられる。軸受中心間距離は約５．５ ｍであり、初段翼長さ約９２ｍｍ，最終段翼長さ約２３５ｍｍである。ダブティルは逆クリ型である。最終段動翼前の静翼に対応するロータシャフトの直径は約６３０ｍｍであり、その直径に対する軸受間距離の比は約８．７ 倍である。
【０１５１】
本実施例の中圧蒸気タービンのロータシャフトは動翼植込み部の軸方向幅が初段から４段，５段及び最終段に従って３段階で段階的に大きくなっており、最終段での幅は初段に対して約１．４ 倍と大きくなっている。
【０１５２】
また、本蒸気タービンのロータシャフトは静翼部に対応した部分の直径が小さくなっており、その幅は初段動翼，２〜３段及び最終段動翼側に従って４段階で段階的に小さくなっており、前者に対する後者の軸方向の幅が約０．７ 倍と小さくなる。
【０１５３】
本実施例においては後述する表５に示す材料を初段ブレード，ノズルに使用される他はＷ，Ｃｏ及びＢを含まない１２％Ｃｒ系鋼が用いられる。本実施例における動翼の翼部の長さは初段から最終段になるに従って各段で長くなっており、蒸気タービンの出力によって初段から最終段までの長さが９０〜３５０ｍｍで、６〜９段で、各段の翼部の長さは下流側が上流側に対して隣り合う長さで１．１ 〜１．２の割合で長くなっている。
【０１５４】
動翼の植込み部は静翼に対応する部分に比較して直径が大きくなっており、その幅は動翼の翼部長さの大きい程その植込み幅は大きくなっている。その幅の動翼の翼部長さに対する比率は初段から最終段で０．５〜０．７であり、初段から最終段になるに従って段階的に小さくなっている。
【０１５５】
図２は低圧タービンの断面図である。低圧タービンは２基タンデムに結合され、ほぼ同じ構造を有している。各々動翼４１は左右に８段あり、左右ほぼ対称になっており、また動翼に対応して静翼４２が設けられる。最終段の動翼長さは ４３インチあり、表３のＮｏ．７の１２％Ｃｒ系鋼が使用され、図３に示すダブルティノン，鞍型ダブティルを有し、ノズルボックス４４は複流型である。ロータシャフト４３はＮｉ３．７５％，Ｃｒ１．７５％，Ｍｏ０．４％，Ｖ０．１５％，Ｃ０．２５％，Ｓｉ０．０５％，Ｍｎ０．１０ ％，残Ｆｅからなるスーパークリーン材の全焼戻しベーナイト組織を有する鍛鋼が用いられる。最終段以外の動翼及び静翼にはいずれもＭｏを０．１％含有する１２％Ｃｒ系鋼が用いられる。内外部ケーシング材にはＣ０．２５ ％の鋳鋼が用いられる。本実施例における軸受４３での中心間距離は７５００ｍｍで、静翼部に対応するロータシャフトの直径は約１２８０ｍｍ，動翼植込み部での直径は２２７５ｍｍである。このロータシャフト直径に対する軸受中心間の距離は約５．９ である。
【０１５６】
図３は１０９２ｍｍ（４３″）長翼の斜視図である。５１は、高速蒸気が突き当たる翼部、５２はロータシャフトへの翼植え込み部、５３は翼の遠心力を支えるためのピンを挿入する穴、５４は蒸気中の水滴によるエロージョンを防止するためのエロージョンシールド（Ｃｏ基合金のステライト板を溶接で接合）である。本実施例における４３″長翼は、エレクトロスラグ再溶解法により溶製し、鍛造熱・処理を行った。
【０１５７】
表３は蒸気タービン用長翼材に係る１２％Ｃｒ系鋼の化学組成（重量％）を示すものである。各試料はそれぞれ１５０ｋｇ真空アーク後エレクトロスラグ溶解し、１１１５０℃に加熱し鍛造して実験素材とした。試料Ｎｏ．１は、１０００℃で１ｈ加熱後油焼入れにより室温まで冷却し、次いで、５７０℃に加熱し２ｈ保持後室温まで空冷した。Ｎｏ．２は、１０５０℃で１ｈ加熱後油焼入れにより室温まで冷却し、次いで、５７０℃に加熱し２ｈ保持後室温まで空冷した。試料Ｎｏ．３〜Ｎｏ．６は、１０５０℃で１ｈ加熱後油焼入れにより室温まで冷却し、次いで、５６０℃に加熱し２ｈ保持後室温まで空冷し（１次焼戻し）、更に５８０℃に加熱し２ｈ保持後室温まで炉冷した（２次焼戻し）。
【０１５８】
表３において、Ｎｏ．３，４及び５は本発明材、Ｎｏ．６は比較材及びＮｏ．１及び２は、現用の２６″長翼材である。
【０１５９】
表４はこれら試料の室温の機械的性質を示す。本発明材（Ｎｏ．３〜５）は、蒸気タービン用長翼材として要求される引張強さ（１２０ｋｇｆ／ｍｍ^２以上又は１２８．５ｋｇｆ／ｍｍ^２以上）及び低温靭性（２０℃Ｖノッチシャルピー衝撃値２．５ｋｇｆ−ｍ／ｃｍ^２以上）を十分満足することが確認された。
【０１６０】
これに対し、比較材のＮｏ．１及び６は、蒸気タービン用長翼に使用するには、引張強さと衝撃値とで示される値が低い。比較材試番２は、引張強さ及び靭性が低い。Ｎｏ．５は、衝撃値が３．８ｋｇｆ−ｍ／ｃｍ^２と若干低く、４３″以上に対しては４ｋｇｆ−ｍ／ｃｍ^２以上の要求に若干不足である。
【０１６１】
【表３】

【０１６２】
【表４】

【０１６３】
本実施例においてはＮｉとＭｏ量とは同等の含有量で含有させることによって低温における強度と靭性とをともに高めるものであり、両者の含有量の差が大きくなるに従って強度が低下する傾向を示す。Ｎｉ量がＭｏ量より０．６ ％以上少なくなると急激に強度が低下し、逆に１．０ ％以上多くなることによっても急激に強度が低下する。従って、（Ｎｉ−Ｍｏ）量が−０．６〜１．０％が高い強度を示す。また、（Ｎｉ−Ｍｏ）量は−０．５％付近で衝撃値が低下するがその前後では高い値を示す。
【０１６４】
焼入れ温度は９７５〜１１２５℃，１次焼戻し５５０〜５６０℃で行った後、２次焼戻し温度は５６０〜５９０℃である。長翼材として要求される特性（引張強さ≧１２８．５ｋｇｆ／ｍｍ^２，２０℃Ｖノッチシャルピー衝撃値≧４ｋｇｆ−ｍ／ ｃｍ^２）を、満足することが確認された。
【０１６５】
本実施例における１２％Ｃｒ系鋼は前述の如く引張強さ１２０ｋｇｆ／ｍｍ^２ 以上及び衝撃値４ｋｇｆ−ｍ／ｃｍ^２以上を有するものが好ましいが、衝撃値（ｙ）が 〔−０．４５×（引張強さ）＋６１．５〕によって求められる値以上とするものが特に好ましいものである。
【０１６６】
本発明に係る１２％Ｃｒ系鋼は特に、Ｃ＋Ｎｂ量が０．１８〜０．３５％で、 （Ｎｂ／Ｃ）比が０．４５〜１．００，（Ｎｂ／Ｎ）比が０．８〜３．０が好ましい。
【０１６７】
本実施例の低圧タービンは動翼植込み部の軸方向の幅が初段〜３段，４段，５段，６〜７段及び８段の４段階で徐々に大きくなっており、最終段の幅は初段の幅に比べ約２．５ 倍と大きくなっている。
【０１６８】
また、静翼部に対応する部分の直径は小さくなっており、その部分の軸方向の幅は初段動翼側から５段目，６段目及び７段目の３段階で徐々に大きくなっており、最終段側の幅は初段側に対して約１．９ 倍大きくなっている。
【０１６９】
本実施例における動翼は８段であり、その翼部長さは初段の３″から５″， ７″，１０″，１３″，１８″，２７″及び４３″の最終段になるに従って各段で長くなっており、蒸気タービンの出力によって初段から最終段の長さが９０〜１２７０ｍｍで、８段又は９段で、各段の翼部長さは下流側が上流側に対して隣り合う長さで１．３〜１．６倍の割合で長くなっている。
【０１７０】
動翼の植込み部は静翼に対応する部分に比較して直径が大きくなっており、その幅は動翼の翼部長さの大きい程その植込み幅は大きくなっている。その幅の動翼の翼部長さに対する比率は初段から最終段で０．１５〜０．９１であり、初段から最終段になるに従って段階的に小さくなっている。
【０１７１】
また、各静翼に対応する部分のロータシャフトの幅は初段と２段目との間から最終段とその手前との間までの各段で段階的に小さくなっている。その幅の動翼の翼部長さに対する比率は０．２５〜１．２５で上流側から下流側になるに従って小さくなっている。
【０１７２】
本実施例の他、高圧蒸気タービン及び中圧蒸気タービンへの蒸気入口温度６１０℃，２基の低圧蒸気タービンへの蒸気入口温度３８５℃とする１０００ＭＷ級大容量発電プラントに対しても同様の構成とすることができる。
【０１７３】
図４は石炭燃焼高温高圧蒸気タービンプラントの代表的なプラント構成図を示すものである。
【０１７４】
本実施例における高温高圧蒸気タービンプラントは主として石炭専焼ボイラ ５１，高圧タービン５２，中圧タービン５３，低圧タービン５４，低圧タービン５５，復水器５６，復水ポンプ５７，低圧給水加熱器系統５８，脱気器５９，昇圧ポンプ６０，給水ポンプ６１，高圧給水加熱器系統６３などより構成されている。すなわち、ボイラ５１で発生した超高温高圧蒸気は高圧タービン５２に入り動力を発生させたのち再びボイラ５１にて再熱されて中圧タービン５３へ入り動力を発生させる。この中圧タービン排気蒸気は、低圧タービン５４，５５に入り動力を発生させた後、復水器５６にて凝縮する。この凝縮液は復水ポンプ５７にて低圧給水加熱器系統５８，脱気器５９へ送られる。この脱気器５９にて脱気された給水は昇圧ポンプ６０，給水ポンプ６１にて高圧給水加熱器６３へ送られ昇温された後、ボイラ５１へ戻る。
【０１７５】
ここで、ボイラ５１において給水は節炭器６４，蒸発器６５，過熱器６６を通って高温高圧の蒸気となる。また一方、蒸気を加熱したボイラ燃焼ガスは節炭器６４を出た後、空気加熱器６７に入り空気を加熱する。ここで、給水ポンプ６１の駆動には中圧タービンからの抽気蒸気にて作動する給水ポンプ駆動用タービンが用いられている。
【０１７６】
このように構成された高温高圧蒸気タービンプラントにおいては、高圧給水加熱器系統６３を出た給水の温度が従来の火力プラントにおける給水温度よりもはるかに高くなっているため、必然的にボイラ５１内の節炭器６４を出た燃焼ガスの温度も従来のボイラに比べてはるかに高くなってくる。このため、このボイラ排ガスからの熱回収をはかりガス温度を低下させないようにする。
【０１７７】
尚、本実施例に代えて同じ高圧タービン，中圧タービン及び１基又は２基の低圧タービンをタンデムに連結し、１台の発電機を回転させて発電するタンデムコンパウンド型発電プラントとしても同様に構成することができる。本実施例の如く、出力１０５０ＭＷ級の発電機においてはその発電機シャフトとしてはより高強度のものが用いられる。特に、Ｃ０．１５〜０．３０％，Ｓｉ０．１〜０．３％，Ｍｎ０．５％以下，Ｎｉ３．２５〜４．５％，Ｃｒ２．０５〜３．０％，Ｍｏ０．２５〜０．６０％，Ｖ０．０５〜０．２０ ％を含有する全焼戻しベーナイト組織を有し、室温引張強さ９３ｋｇｆ／ｍｍ^２以上、特に１００ｋｇｆ／ｍｍ^２以上，５０％ＦＡＴＴが０℃以下、特に−２０℃以下とするものが好ましく、２１．２ＫＧ における磁化力が９８５ＡＴ／ｃｍ以下とするもの、不純物としてのＰ，Ｓ，Ｓｎ，Ｓｂ， Ａｓの総量を０．０２５％以下，Ｎｉ／Ｃｒ比を２．０以下とするものが好ましい。
【０１７８】
図５は高圧及び図６は中圧タービンロータシャフトの正面図である。図５の高圧タービンシャフトは多段側の初段ブレード植設部を中心に８段のブレードが植設される構造である。中圧タービンシャフトは多段ブレードが左右に各６段ほぼ対称にブレード植設部が設けられ、ほぼ中心を境にしたものである。低圧タービン用ロータシャフトは図示されていないが、高圧，中圧，低圧タービンのいずれのロータシャフトにおいても中心孔が設けられ、この中心孔を通して超音波検査，目視検査及びけい光探傷によって欠陥の有無が検査される。
【０１７９】
表５は本実施例の高圧タービン，中圧タービン及び低圧タービンの主要部に用いた化学組成（重量％）を示す。本実施例においては、高圧部及び中圧部の高温部を全部フェライト系の結晶構造を有する熱膨張係数１２×１０^−６／℃のものにしたので、熱膨張係数の違いによる問題は全くなかった。
【０１８０】
高圧部及び中圧部のロータは、表５に記載の胴部及び軸受部に係る耐熱鋳鋼を電気炉で３０トン溶解し、カーボン真空脱酸し、金型鋳型に鋳込み，鍛伸して電極棒を作製し、この電極棒を用い、先ず軸受部をエレクトロスラグ溶解した後、直ちに胴部についてエレクトロ再溶解し、更に軸受部をその上にエレクトロスラグ再溶解し、ロータ形状（直径１０５０ｍｍ，長さ３７００ｍｍ）に鍛伸して成型した。この鍛伸は、鍛造割れを防ぐために、１１５０℃以下の温度で行った。またこの鍛鋼を焼鈍熱処理後、１０５０℃に加熱し水噴霧冷却焼入れ処理、５７０℃及び６９０℃で２回焼戻しを行い、図５及び図６に示す形状に切削加工によって得たものである。胴部と軸受部とは点線に示す位置で接合したものである。図５に示すように高圧蒸気タービン用ロータシャフトではブレードの下流側最終段とその手前との間、図６に示す中圧蒸気タービン用ロータシャフトでは下流側最終段とその手前との間で各々接合したものである。本実施例においてはエレクトロスラグ鋼塊の上部側を胴部の初段翼側にし、下部を最終段側にするようにした。高圧部及び中圧部のブレード及びノズルは、同じく表５に記載の耐熱鋼を真空アーク溶解炉で溶解し、ブレード及びノズル素材形状（幅１５０ｍｍ，高さ５０ｍｍ，長さ１０００ｍｍ）に鍛伸して成型した。この鍛伸は、鍛造割れを防ぐために、１１５０℃以下の温度で行った。またこの鍛鋼を１０５０℃に加熱し油焼入れ処理し、６９０℃で焼戻しを行い、次いで所定形状に切削加工したものである。
【０１８１】
高圧部及び中圧部の内部ケーシング，主蒸気止め弁ケーシング及び蒸気加減弁ケーシングは、表５に記載の耐熱鋳鋼を電気炉で溶解し、とりべ精錬後、砂型鋳型に鋳込み作製した。鋳込み前に、十分な精錬及び脱酸を行うことにより、引け巣等の鋳造欠陥のないものができた。このケーシング材を用いた溶接性評価は、ＪＩＳ Ｚ３１５８に準じて行った。予熱，パス間及び後熱開始温度は２００℃に、後熱処理は４００℃×３０分にした。本発明材には溶接割れが認められず、溶接性が良好であった。
【０１８２】
【表５】

【０１８３】
表６は、上述したフェライト系鋼製高温蒸気タービン主要部材を切断調査した機械的性質及び熱処理条件を示す。
【０１８４】
このロータシャフトの中心部を調査した結果、高圧，中圧タービンロータに要求される特性（６２５℃，１０^５ｈ強度≧１３ｋｇｆ／ｍｍ^２，２０℃衝撃吸収エネルギー≧１．５ｋｇｆ−ｍ）を十分満足することが確認された。これにより、６２０℃以上の蒸気中で使用可能な蒸気タービンロータが製造できることが実証された。またこのブレードの特性を調査した結果、高圧，中圧タービンの初段ブレードに要求される特性（６２５℃，１０^５ｈ強度≧１５ｋｇｆ／ｍｍ^２）を十分満足することが確認された。これにより、６２０℃以上の蒸気中で使用可能な蒸気タービンブレードが製造できることが実証された。
【０１８５】
さらにこのケーシングの特性を調査した結果、高圧，中圧タービンケーシングに要求される特性（６２５℃，１０^５ｈ強度≧１０ｋｇｆ／ｍｍ^２，２０℃衝撃吸収エネルギー≧１ｋｇｆ−ｍ）を十分満足することと溶接可能であることが確認された。これにより、６２０℃以上の蒸気中で使用可能な蒸気タービンケーシングが製造できることが実証された。
【０１８６】
【表６】

【０１８７】
本実施例においては、ロータシャフトのジャーナル部にＣｒ−Ｍｏ低合金鋼を肉盛溶接し、軸受特性を改善させた。肉盛溶接は次の通りである。
【０１８８】
供試溶接棒として被覆アーク溶接棒（直径４．０φ）を用いた。その溶接棒を用いて溶接したものの溶着金属の化学組成（重量％）を表７に示す。この溶着金属の組成は溶接材の組成とほぼ同じである。
【０１８９】
溶接条件は溶接電流１７０Ａ，電圧２４Ｖ，速度２６ｃｍ／ｍｉｎ である。
【０１９０】
【表７】

【０１９１】
肉盛溶接を上述の供試母材表面に表８に示すごとく、Ｎｏ．１及びＮｏ．２の２種について各層ごとに使用溶接棒を組合せて、８層の溶接を行った。各層の厚さは３〜４ｍｍであり、全厚さは約２８ｍｍであり、表面を約５ｍｍ研削した。
【０１９２】
溶接施工条件は、予熱，パス間，応力除去焼鈍（ＳＲ）開始温度が２５０〜３５０℃及びＳＲ処理条件は６３０℃×３６時間保持である。
【０１９３】
【表８】

【０１９４】
溶接部の性能を確認するために板材に同様に肉盛溶接し、１６０゜の側曲げ試験を行ったが、いずれも溶接部に割れは認められなかった。いずれも６層目以降が各々の表に示す組成を有するものである。
【０１９５】
更に、本発明における回転による軸受摺動試験を行ったが、いずれも軸受に対する悪影響もなく、耐酸化性に対しても優れたものであった。
【０１９６】
本実施例に代えて高圧蒸気タービン，中圧蒸気タービン及び２基の低圧蒸気タービンをタンデムに結合し、３６００回転としたタンデム型発電プラントにおいても同様に構成できるものである。
【０１９７】
（実施例２）
表９は蒸気温度６２１℃，６００ＭＷ蒸気タービンの主な仕様である。本実施例は、タンデムコンパウンドダブルフロー型，低圧タービンにおける最終段翼長が４３インチであり、ＨＰ・ＩＰ一体型及びＬＰ１台で３０００ｒｐｍ／ｍｉｎの回転数を有し、高温部においては表に示す主な材料によって構成される。高圧部（ＨＰ）の蒸気温度は６００℃，２５０ｋｇ／ｃｍ^２ の圧力であり、中圧部（ＩＰ）の蒸気温度は６００℃に再熱器によって加熱され、１７０〜１８０ｋｇ／ｃｍ^２ の圧力で運転される。低圧部（ＬＰ）は蒸気温度は４５０℃で入り、１００℃以下，７２２ｍｍＨｇの真空で復水器に送られる。
【０１９８】
【表９】

【０１９９】
図７は高圧中圧一体型蒸気タービンの断面構成図及び図８はそのロータシャフトの断面図である。高圧側蒸気タービンは内部車室１８とその外側の外部車室 １９内に高圧側動翼１６を植設した高中圧車軸（高圧ロータシャフト）２３が設けられる。前述の高温高圧の蒸気は前述のボイラによって得られ、主蒸気管を通って、主蒸気入口を構成するフランジ，エルボ２５より主蒸気入口２８を通り、ノズルボックス３８より初段の動翼に導かれる。動翼は図中左側の高圧側に８段及び（図中右側約半分の）中圧側に６段設けられる。これらの動翼に対応して各々静翼が設けられる。動翼は鞍型ダブティル型式，ダブルティノン，高圧側初段翼長約４０ｍｍ，中圧側初段翼長が１３０ｍｍである。軸受４３間の長さは約５．７ｍ及び静翼部に対応する部分で最も小さい部分の直径は約７４０ｍｍであり、直径に対する長さの比は約７．７ である。高中圧一体ロータシャフトにおいても中心孔が設けられ、欠陥の有無が検査される。
【０２００】
高圧側ロータシャフトの初段と最終段の動翼植込み付根部分の幅は初段が最も広く、２段目〜７段目がそれより小さく、初段の０．４０〜０．５６倍でいずれも同等の大きさであり、最終段が初段と２〜７段目の大きさの間にあり、初段の ０．４６〜０．６２倍の大きさである。
【０２０１】
高圧側においてはブレード及びノズルを後述する表５に示す１２％Ｃｒ系鋼によって構成したものである。本実施例における動翼の翼部の長さは初段が３５〜５０ｍｍ、２段目から最終段になるに従って各段で長くなっており、特に蒸気タービンの出力によって２段から最終段までの長さが５０〜１５０ｍｍの範囲内であり、段数は７〜１２段の範囲内にあり、各段の翼部の長さは下流側が上流側に対して隣り合う長さで１．０５〜１．３５倍の範囲内で長くなっているとともに、下流側でその比率が徐々に大きくなっている。
【０２０２】
中圧側蒸気タービンは高圧側蒸気タービンより排出された蒸気を再度６００℃に再熱器によって加熱された蒸気によって高圧蒸気タービンと共に発電機を回転させるもので、３０００ｒｐｍ の回転数によって回転される。中圧側タービンは高圧側タービンと同様に内部車室２１と外部車室２２とを有し、動翼１７と対抗して静翼が設けられる。動翼１７は６段である。初段翼長さ約１３０ｍｍ，最終段翼長さ約２６０ｍｍである。ダブティルは逆クリ型である。静翼に対応するロータシャフトの直径は約７４０ｍｍである。
【０２０３】
高中圧蒸気タービンのロータシャフトは動翼植込み付根部の軸方向幅が初段が最も大きく、２段目がそれより小さく、３〜５段目が２段目より小さくいずれも同じで、最終段の幅は３〜５段目と２段目の間の大きさで、初段の０．４８〜 ０．６４倍である。初段は２段目の１．１〜１．５倍である。
【０２０４】
中圧側においてはブレード及びノズルを後述する表５に示す１２％Ｃｒ系鋼が用いられる。本実施例における動翼の翼部の長さは初段から最終段になるに従って各段で長くなっており、蒸気タービンの出力によって初段から最終段までの長さが９０〜３５０ｍｍ，段数が６〜９段の範囲内にあり、各段の翼部の長さは下流側が上流側に対して隣り合う長さで１．１０〜１．２５の割合で長くなっている。動翼の植込み部は静翼に対応する部分に比較して直径が大きくなっており、その幅は動翼の翼部長さと位置に関係する。その幅の動翼の翼部長さに対する比率は初段が最も大きく、１．３５〜１．８０倍，２段目が０．８８〜１．１８倍，３〜６段目が最終段になるに従って小さくなっており、０．４０〜０．６５倍である。図９は低圧タービンの断面図及び図１０はそのロータシャフトの断面図である。低圧タービンは１基で高中圧にタンデムに結合される。動翼４１は左右に６段あり、左右ほぼ対称になっており、また動翼に対応して静翼４２が設けられる。最終段の動翼長さは４３インチあり、表３に示す１２％Ｃｒ系鋼又はＴｉ基合金が使用され、図３に示すいずれもダブルティノン，鞍型ダブティルを有し、ノズルボックス４４は複流型である。Ｔｉ基合金は時効硬化処理が施され、重量で Ａｌ６％，Ｖ４％を含むものである。ロータシャフト４３はＮｉ３．７５ ％， Ｃｒ１．７５％，Ｍｏ０．４％，Ｖ０．１５％，Ｃ０．２５％，Ｓｉ０．０５ ％， Ｍｎ０．１０ ％，残Ｆｅからなるスーパークリーン材の全焼戻しベーナイト組織を有する鍛鋼が用いられる。最終段以外の動翼及び静翼にはいずれもＭｏを０．１％含有する１２％Ｃｒ系鋼が用いられる。内外部ケーシング材にはＣ０．２５ ％の鋳鋼が用いられる。本実施例における軸受４３での中心間距離は６０００ｍｍで、静翼部に対応するロータシャフトの直径は約８９０ｍｍ、動翼植込み部での直径は各段同じで１８２０ｍｍである。静翼部に対応するロータシャフト直径に対する軸受中心間の距離は約６．７ である。
【０２０５】
低圧タービンは動翼植込み付根部の軸方向の幅が初段が最も小さく、下流側に従って２，３段が同等、４段，５段が同等で４段階で徐々に大きくなっており、最終段の幅は初段の幅に比べ３．８〜４．８倍と大きくなっている。２，３段は初段の１．０５〜１．４０倍、４，５段が２，３段の１．０５〜１．３５倍、最終段が４，５段の２．８〜３．２倍となっている。付根部の幅は末広がりの延長線とロータシャフトの直径とを結ぶ点で示す。
【０２０６】
本実施例における動翼の翼部長さは初段の５″から７″，９″，１６″，２６″及び４３″の最終段になるに従って各段で長くなっており、蒸気タービンの出力によって初段から最終段の長さが８０〜１２７０ｍｍの範囲内で、最大で９段で、各段の翼部長さは下流側が上流側に対して隣り合う長さで１．３〜１．９倍の範囲内で長くなっている。
【０２０７】
動翼の植込み付根部は静翼に対応する部分に比較して直径が大きく末広がりになっており、その幅は動翼の翼部長さの大きい程その植込み幅は大きくなっている。その幅の動翼の翼部長さに対する比率は初段から最終段を除き０．３０〜１．５であり、初段から最終段になるに従って徐々に小さくなっており、後段の比率はその手前のものより０．１５〜０．４０の範囲内で小さくなっている。最終段は ０．５５〜０．６５の比率である。
【０２０８】
本実施例における最終段動翼は実施例１と同じである。
【０２０９】
本実施例の他、高中圧蒸気タービンの蒸気入口温度６１０℃以上，低圧蒸気タービンへの蒸気入口温度約４５０℃及び出口温度が約６０℃とする１０００ＭＷ級大容量発電プラントに対しても同様の構成とすることができる。
【０２１０】
本実施例における高温高圧蒸気タービン発電プラントは主としてボイラ，高中圧タービン，低圧タービン，復水器，復水ポンプ，低圧給水加熱器系統，脱気器，昇圧ポンプ，給水ポンプ，高圧給水加熱器系統などより構成される。すなわち、ボイラで発生した超高温高圧蒸気は高圧側タービンに入り動力を発生させたのち再びボイラにて再熱されて中圧側タービンへ入り動力を発生させる。この高中圧タービン排気蒸気は、低圧タービンに入り動力を発生させた後、復水器にて凝縮する。この凝縮液は復水ポンプにて低圧給水加熱器系統，脱気器へ送られる。この脱気器にて脱気された給水は昇圧ポンプ，給水ポンプにて高圧給水加熱器へ送られ昇温された後、ボイラへ戻る。
【０２１１】
ここで、ボイラにおいて給水は節炭器，蒸発器，過熱器を通って高温高圧の蒸気となる。また一方、蒸気を加熱したボイラ燃焼ガスは節炭器を出た後、空気加熱器に入り空気を加熱する。ここで、給水ポンプの駆動には中圧タービンからの抽気蒸気にて作動する給水ポンプ駆動用タービンが用いられている。
【０２１２】
このように構成された高温高圧蒸気タービンプラントにおいては、高圧給水加熱器系統６３を出た給水の温度が従来の火力プラントにおける給水温度よりもはるかに高くなっているため、必然的にボイラ内の節炭器を出た燃焼ガスの温度も従来のボイラに比べてはるかに高くなってくる。このため、このボイラ排ガスからの熱回収をはかりガス温度を低下させないようにする。
【０２１３】
尚、本実施例では高中圧タービン及び１基の低圧タービンを１台の発電機タンデムに連結し発電するタンデムコンパウンドダブルフロー型発電プラントに構成したものである。別の実施例として、出力１０５０ＭＷ級の発電機においてはその発電機シャフトとしてはより高強度のものが用いられる。特に、Ｃ０．１５ 〜０．３０％，Ｓｉ０．１〜０．３％，Ｍｎ０．５％以下，Ｎｉ３．２５〜４．５％， Ｃｒ２．０５〜３．０％，Ｍｏ０．２５〜０．６０％，Ｖ０．０５〜０．２０％を含有する全焼戻しベーナイト組織を有し、室温引張強さ９３ｋｇｆ／ｍｍ^２ 以上、特に １００ｋｇｆ／ｍｍ^２ 以上，５０％ＦＡＴＴが０℃以下、特に−２０℃以下とするものが好ましく、２１．２ＫＧ における磁化力が９８５ＡＴ／ｃｍ以下とするもの、不純物としてのＰ，Ｓ，Ｓｎ，Ｓｂ，Ａｓの総量を０．０２５ ％以下，Ｎｉ／ Ｃｒ比を２．０ 以下とするものが好ましい。
【０２１４】
前述の表５は本実施例の高中圧タービン及び低圧タービンの主要部に用いた化学組成（重量％）を示す。本実施例においては、高圧側及び中圧側とを一体にした高温部後述の実施例４のＮｏ．９のマルテンサイト鋼を使用した他は表５のものを用い、全部フェライト系の結晶構造を有する熱膨張係数１２×１０^−６／℃のものにしたので、熱膨張係数の違いによる問題は全くなかった。
【０２１５】
高中圧蒸気タービン用ロータシャフト胴部は、表１０に示す１２％Ｃｒ系耐熱鋼が用いられ、本実施例においてはＮｏ．２に記載の耐熱鋳鋼を用いて、実施例１と同様に胴部及び軸受部の電極の製造及び電極からのロータシャフトに相当する鋼塊を製造するとともに、ロータ形状（直径１４５０ｍｍ，長さ５０００ｍｍ）に鍛伸して成型した。この鍛伸は、鍛造割れを防ぐために、１１５０℃以下の温度で行った。またこの鍛鋼を焼鈍熱処理後、１０５０℃に加熱し水噴霧冷却焼入れ処理、５７０℃及び６９０℃で２回焼戻しを行い、図８に示す形状に切削加工によって得たものである。図に示す点線部で胴部と軸受部とを別々の材料で製造した。更に、軸受ジャーナル部４５への肉盛溶接も同様に行った。
【０２１６】
（実施例３）
表１０に示す組成の合金を真空溶解によって、１０ｋｇのインゴットに鋳造し、３０ｍｍ角に鍛造したものである。大型蒸気タービンロータシャフトの場合には、その中心部を模擬して１０５０℃×５時間１００℃／ｈ冷却の焼入れ、５７０℃×２０時間の１次焼戻しと６９０℃×２０時間の２次焼戻し及びブレードにおいては１１００℃×１時間の焼入れ、７５０℃×１時間の焼戻しを行って、６２５℃，３０ｋｇｆ／ｍｍ^２ でクリープ破断試験を実施した。結果を表７に合わせて示す。
【０２１７】
表１０のＮｏ．１〜Ｎｏ．６の本発明合金は、６２０℃以上の蒸気条件に適用するのに好ましいもので、クリープ破断寿命が長いことがわかる。Ｃｏ量が多い程クリープ破断時間が向上するが、Ｃｏの多量の増加は６００〜６６０℃で加熱を受けると加熱脆化が生じる傾向を有するので、強化と靭性の両方を高めるには６２０〜６３０℃に対しては２〜５％、６３０〜６６０℃に対しては５．５〜８ ％が好ましい。Ｂは０．０３ ％以下が優れた強度を示す。６２０〜６３０℃ではＢ量を０．００１〜０．０１％及びＣｏ量を２〜４％、６３０〜６６０℃のより高温側ではＢ量を０．０１〜０．０３％とし、Ｃｏ量を５〜７．５ ％と高めることにより高強度が得られる。
【０２１８】
Ｎは本実施例における６００℃を越える温度では少ない方が強化され、Ｎ量の多いものに比べて強度が高いことが明らかとなった。Ｎ量は０．０１〜０．０４％が好ましい。真空溶解においてはＮはほとんど含有されないので、母合金によって添加したものである。
【０２１９】
表１０に示すように、ロータ材は本実施例のＮｏ．２の合金に相当し、高い強度が得られる。Ｍｎ量を０．０９ ％と低くすると同じＣｏ量で比較して高い強度を示すことから、より強化のためにはＭｎ量を０．０３〜０．２０％とするのが好ましい。
【０２２０】
【表１０】

【０２２１】
（実施例４）
表１１は本発明の高圧，中圧及び高中圧タービン用内部ケーシング材に係る化学組成（重量％）を示す。試料は大型ケーシングの厚肉部を想定して、高周波誘導溶解炉を用い２００ｋｇ溶解し、最大厚さ２００ｍｍ，幅３８０ｍｍ，高さ４４０ｍｍの砂型に鋳込み，鋳塊を作製した。試料は、１０５０℃×８ｈ炉冷の焼鈍処理後、大型蒸気タービンケーシングの厚肉部を想定して焼準（１０５０℃×８ｈ→空冷），焼戻し（７１０℃×７ｈ→空冷，７１０℃×７ｈ→空冷の２回）の熱処理を行った。
【０２２２】
溶接性評価は、ＪＩＳ Ｚ３１５８に準じて行った。予熱，パス間及び後熱開始温度は１５０℃に、後熱処理は４００℃×３０分にした。
【０２２３】
【表１１】

【０２２４】
表１２は室温の引張特性、２０℃におけるＶノッチシャルピー衝撃吸収エネルギー、６５０℃，１０^５ ｈクリープ破断強度及び溶接割れ試験結果を示す。
【０２２５】
適量のＢ，Ｍｏ及びＷを添加した本発明材のクリープ破断強度及び衝撃吸収エネルギーは、高温高圧タービンケーシングに要求される特性（６２５℃，１０^５ｈ強度≧８ｋｇｆ／ｍｍ^２，２０℃衝撃吸収エネルギー≧１ｋｇｆ−ｍ）を十分満足する。特に、９ｋｇｆ／ｍｍ^２ 以上の高い値を示している。また、本発明材には溶接割れが認められず、溶接性が良好である。Ｂ量と溶接割れの関係を調べた結果、Ｂ量が０．００３５ ％を越えると、溶接割れが発生した。Ｎｏ．１のものは若干割れの心配があった。機械的性質に及ぼすＭｏの影響を見ると、Ｍｏ量を１．１８ ％と多いものは、クリープ破断強度は高いものの、衝撃値が低く、要求される靭性を満足できなかった。一方、Ｍｏ０．１１ ％のものは、靭性は高いものの、クリープ破断強度が低く、要求される強度を満足できなかった。
【０２２６】
機械的性質に及ぼすＷの影響を調べた結果、Ｗ量を１．１ ％以上にするとクリープ破断強度が顕著に高くなるが、逆にＷ量を２％以上にすると室温衝撃吸収エネルギーが低くなる。特に、Ｎｉ／Ｗ比を０．２５〜０．７５に調整することにより、温度６２１℃，圧力２５０ｋｇｆ／ｃｍ^２ 以上の高温高圧タービンの高圧及び中圧内部ケーシング並びに主蒸気止め弁及び加減弁ケーシングに要求される、６２５℃，１０^５ｈクリープ破断強度９ｋｇｆ／ｍｍ^２以上，室温衝撃吸収エネルギー１ｋｇｆ−ｍ以上の耐熱鋳鋼ケーシング材が得られる。特に、Ｗ量１．２ 〜２％，Ｎｉ／Ｗ比を０．２５〜０．７５ に調整することにより、６２５℃，１０^５ｈクリープ破断強度１０ｋｇｆ／ｍｍ^２ 以上，室温衝撃吸収エネルギー２ｋｇｆ−ｍ以上の優れた耐熱鋳鋼ケーシング材が得られる。
【０２２７】
【表１２】

【０２２８】
Ｗ量は１．０％以上とすることによって顕著に強化されるとともに、特に１．５％以上では８．０ｋｇｆ／ｍｍ^２以上の値が得られる。本発明のＮｏ．７は６４０℃以下で十分要求の強度を満足するものであった。
【０２２９】
本発明の耐熱鋳鋼を目標組成とする合金原料を電気炉で１トン溶解し、とりべ精錬後、砂型鋳型に鋳込み実施例３に記載の高中圧部の内部ケーシングを得た。このケーシングを１０５０℃×８ｈ炉冷の焼鈍熱処理後、１０５０℃×８ｈ衝風冷の焼準熱処理，７３０℃×８ｈ炉冷の２回焼戻しを行った。全焼戻しマルテンサイト組織を有するこの試作ケーシングを切断調査した結果、２５０気圧，６２５℃高温高圧タービンケーシングに要求される特性（６２５℃，１０^５ｈ強度≧９ ｋｇｆ／ｍｍ^２ ，２０℃衝撃吸収エネルギー≧１ｋｇｆ−ｍ）を十分満足することと溶接可能であることが確認できた。
【０２３０】
（実施例５）
本実施例においては、高圧蒸気タービン及び中圧蒸気タービン又は高中圧蒸気タービンの蒸気温度を６２５℃に代えて６４９℃としたものであり、構造及び大きさを実施例２又は３とほぼ同じ設計で得られるものである。ここで実施例１と変わるものはこの温度に直接接する高圧，中圧又は高中圧蒸気タービンのロータシャフト，初段動翼及び初段静翼と内部ケーシングである。内部ケーシングを除くこれらの材料としては前述の表７に示す材料のうちＢ量を０．０１〜０．０３％及びＣｏ量を５〜７％と高め、更に内部ケーシング材としては実施例１のＷ量を２〜３％に高め、Ｃｏを３％加えることにより、要求される強度が満足し、従来の設計が使用できる大きなメリットがある。即ち、本実施例においては高温にさらされる構造材料が全てフェライト系鋼によって構成される点に従来の設計思想がそのまま使用できるのである。尚、２段目の動翼及び静翼の蒸気入口温度は約６１０℃となるので、これらには実施例１の初段に用いた材料を用いることが好ましい。
【０２３１】
更に、低圧蒸気タービンの蒸気温度は実施例２又は３の約３８０℃に比べ若干高い約４０５℃となるが、そのロータシャフト自身は実施例１の材料が十分に高強度を有するので、同じくスーパークリーン材が用いられる。
【０２３２】
更に、本実施例におけるクロスコンパウンド型に対し、全部を直結したタンデム型で３６００ｒｐｍ の回転数においても実施できるものである。
【０２３３】
【発明の効果】
発明によれば、より高合金化されたマルテンサイト鋼からなるロータシャフトにおいて高い溶接性を有するジャーナル部を有する蒸気タービン用ロータシャフトが得られる。
【図面の簡単な説明】
【図１】本発明に係る高圧，中圧蒸気タービンの断面図。
【図２】本発明に係る低圧蒸気タービンの断面構造図。
【図３】本発明に係るタービン動翼の斜視図。
【図４】本発明に係る石炭燃焼発電プラントの構成図。
【図５】本発明に係る高圧蒸気タービン用ロータシャフトの断面図。
【図６】本発明に係る中圧蒸気タービン用ロータシャフトの断面図。
【図７】本発明に係る高中圧蒸気タービンの断面図。
【図８】本発明に係る高中圧蒸気タービン用ロータシャフトの断面図。
【図９】本発明に係る低圧蒸気タービンの断面図。
【図１０】本発明に係る低圧蒸気タービン用ロータシャフトの断面図。
【符号の説明】
１…第１軸受、２…第２軸受、３…第３軸受、４…第４軸受、５…推力軸受、１０…第１シャフトパッキン、１１…第２シャフトパッキン、１２…第３シャフトパッキン、１３…第４シャフトパッキン、１４…高圧隔板、１５…中圧隔板、１６…高圧動翼、１７…中圧動翼、１８…高圧内部車室、１９…高圧外部車室、２０…中圧内部第１車室、２１…中圧内部第２車室、２２…中圧外部車室、２３…高圧車軸、２４…中圧車軸、２５…フランジ，エルボ、２６…前側軸受箱、 ２７…ジャーナル部、２８…主蒸気入口、２９…再熱蒸気入口、３０…高圧蒸気排気口、３１…気筒連絡管、３８…ノズルボックス（高圧第１段）、３９…推力軸受摩耗遮断装置、４０…暖機蒸気入口、４３…軸受、５１…ボイラ、５２…高圧タービン、５３…中圧タービン、５４，５５…低圧タービン、５６…復水器、５７…復水ポンプ、５８…低圧給水加熱器系統、５９…脱気器、６０…昇圧ポンプ、６１…給水ポンプ、６３…高圧給水加熱器系統、６４…節炭器、６５…蒸発器、６６…過熱器、６７…空気加熱器、６８…発電機。[0001]
BACKGROUND OF THE INVENTION
The present invention relates to a novel rotor shaft for a steam turbine and a method for manufacturing the same, and in particular, a rotor shaft for an ultra super critical pressure steam turbine in which a journal portion is made of martensite steel with good weldability and a body portion is made of martensite steel having high high-temperature strength. And a manufacturing method thereof, a steam turbine power plant and a steam turbine thereof.
[0002]
[Prior art]
The conventional steam turbine has a maximum steam temperature of 566 ° C. and a steam pressure of 246 atg. As this rotor material, 1Cr-1Mo-1 / 4V low alloy steel or 11Cr-1Mo-V-Nb-N steel shown in Japanese Patent Publication No. 40-4137 is used.
[0003]
However, from the viewpoint of depletion of fossil fuels such as oil and coal and energy saving, higher efficiency of thermal power plants is desired. In order to increase the power generation efficiency, the most effective means is to increase the steam temperature of the steam turbine. As these high-efficiency turbine materials, the current rotor material is insufficient in strength, and a material having higher strength than this is required.
[0004]
However, since all of the above-mentioned alloys have insufficient high-temperature strength as a high-temperature steam turbine rotor having a steam temperature of 621 ° C. or higher, the inventors have disclosed 11Cr shown in JP-A-4-147948 having high high-temperature strength. -W-Co-Mo-V-Nb-NB steel is being developed. However, although 12% Cr steel has excellent mechanical properties, it has extremely poor wear resistance. Therefore, a damage accident occurs between the rotor journal portion and the bearing metal. This damage to the journal is considered to be caused by foreign material intrusion between the journal and the bearing metal. In particular, 12% Cr heat-resisting steel has a low thermal conductivity and a high Cr content. Therefore, it is considered that Cr carbide is easily generated, and promotes damage to the journal portion. In order to prevent this 12% Cr heat resistant steel rotor journal from being damaged, overlay welding of a low alloy steel having excellent wear resistance is the most excellent method. For this reason, the above 12% Cr steel rotor material has high strength at high temperatures and the journal portion has poor slidability. Therefore, JP-A-57-105502 discloses an integrated rotor shaft in which the journal portion is made of low alloy steel. Has been.
[0005]
[Problems to be solved by the invention]
In the invention described in the above publication, when the temperature of the steam turbine becomes higher, the temperature at the journal portion becomes higher at the same time as the strength at the barrel portion becomes higher, so there is a problem in strength in the journal portion made of low alloy steel. There is. Furthermore, although build-up welding of low alloy steel to the journal part is known, the inventors have found that build-up welding for high alloying consisting of large structures as described above is extremely difficult, and the present invention. It came to.
[0006]
Furthermore, from the viewpoint of depletion of fossil fuels such as oil and coal, energy saving, and prevention of environmental pollution, higher efficiency of thermal power plants is desired. In order to increase the power generation efficiency, the most effective means is to increase the steam temperature of the steam turbine. JP-A-7-233704 is known as a material for these high-efficiency ultra-high temperature steam turbines.
[0007]
The present invention has been made in order to cope with the recent increase in the length of low-pressure steam turbine blades. Japanese Patent Application Laid-Open Nos. 63-171856 and 4-120246 completely disclose the blade material for a steam turbine. Not disclosed.
[0008]
In addition, in the publication described above in Japanese Patent Laid-Open No. 7-233704, the rotor material and the casing material are disclosed. As described above, the final stage moving blade in the high-medium pressure integrated steam turbine and the low-pressure steam turbine at higher temperatures. There is no description about 12% Cr martensitic steel.
[0009]
An object of the present invention is to provide a steam turbine, a steam turbine power plant, and a rotor shaft having a journal portion having a high weldability in a rotor shaft made of a highly alloyed martensitic steel.
[0011]
[Means for Solving the Problems]
The steam turbine rotor shaft of the present invention is composed of a 12% Cr alloy steel having a good weldability in the journal part and the low temperature region part, and a body part integrally made of 12% Cr alloy steel having a high temperature strength higher than that of the journal part. It is.
[0012]
In the rotor shaft for an ultra-supercritical turbine of the present invention, the journal part has a weight ratio of C 0.06 to 0.14%, Si 0.5% or less, Mn 2% or less, Cr 7 to13%, Ni 0.2 to 2%, V 0.05 to 0.35%, Nb 0.01 to 0.20%, N 0.005 to 0.05%, Mo 1% or less, W 3% or less, B added or 0.00%. It is made of martensitic steel containing 0.003% or less and Co5% or less, and the body portion is C0.06-0.14% by weight, Si 0.15% or less, Mn 0.03-1.5%, Cr9-13%, Ni 0.05-1.0%, V 0.05-0 . 3%, Nb 0.01 to 0.20%, N 0.0050 . 05%, Mo 0.05-0.5%, W 1.0-3.5%, B0 . 001It is preferably made of martensitic steel containing 0.03% and Co 10% or less, and is made of 12% Cr alloy steel having higher high-temperature strength than the journal part.
[0013]
The rotor shaft for an ultra-supercritical pressure turbine according to the present invention separately prepares two or more consumable electrodes of alloy steel having a journal portion having higher weldability than the barrel portion or a barrel portion having higher high-temperature strength than the journal portion. First, the consumable electrode corresponding to the former journal part is electroslag-melted, and as soon as a desired length is obtained, the consumable electrode corresponding to the latter body part is electroslag-melted and joined, and then the former journal again. The consumable electrode corresponding to the part can be manufactured by electroslag melting, adding and joining together.
[0014]
In addition, the rotor shaft for an ultra super critical pressure turbine of the present invention is manufactured with alloy steel (upper end and lower end) having good weldability in the journal portion and the low-temperature region portion and alloy steel having high high-temperature strength at the body portion (center portion). Such a consumable electrode can be prepared, and the consumable electrode can be manufactured by electroslag melting.
[0015]
Furthermore, a build-up weld layer having a desired thickness is formed on the bearing portion of the rotor shaft for an ultra super critical pressure turbine according to the present invention with a low alloy steel having high bearing characteristics.
[0016]
The present invention relates to a steam turbine power plant comprising a low pressure turbine or a high / medium pressure integrated steam turbine, and one low pressure turbine connected to one or two tandems, wherein a high pressure turbine and an intermediate pressure turbine are connected. The high-pressure turbine and the medium-pressure turbine or the high-medium-pressure turbine have a steam inlet temperature to the first stage blade of 600 to 660 ° C. (preferably 600 to 620 ° C., 620 to 630 ° C., 630 to 640 ° C.). The rotor shaft or rotor exposed to the steam inlet temperature of the high-pressure turbine and the intermediate-pressure turbine or the high-medium-pressure turbine for the range of the steam inlet temperature to the first stage blades of 380 to 475 ° C. (preferably 400 to 430 ° C.) The shaft, rotor blade, stationary blade and inner casing are made of high strength martensitic steel containing 8 to 13 wt% Cr. The bearing portion of the rotor shaft is more weldable than the body portion, and more preferably, the value of [blade length (inch) × rotational speed (rpm)] of the last stage rotor blade of the low-pressure turbine is 125, It is in the steam turbine power plant characterized by being 000 or more.
[0017]
The present invention further includes a rotor shaft, a moving blade implanted in the rotor shaft, a stationary blade that guides the inflow of water vapor into the moving blade, and an inner casing that holds the stationary blade, The temperature flowing into the first stage of the rotor blade is 600 to 660 ° C. and the pressure is 250 kg / cm.²Or more (preferably 246 to 316 kg / cm²) Or 170-200 kg / cm²The rotor shaft, or the rotor shaft and at least the first stage of the rotor blade and the stationary blade correspond to each steam temperature (preferably 610 ° C, 625 ° C, 640 ° C, 650 ° C, 660 ° C). 10 at temperature⁵Time creep rupture strength is 10kgf / mm²Or more (preferably 12 kgf / mm²Or higher) martensitic steel having a total tempered martensitic structure containing 9.5 to 13 wt% (preferably 10.5 to 11.5 wt%) of Cr, and the bearing portion of the rotor shaft is the same as that described above. Preferably, the inner casing is 10 at a temperature corresponding to each steam temperature.⁵Time creep rupture strength is 10kgf / mm²Or more (preferably 10.5 kgf / mm²The steam emitted from the high-pressure, intermediate-pressure steam turbine or high-pressure side turbine is characterized by being made of martensitic cast steel containing 8 to 9.5% by weight of Cr, which is equal to or higher than the high-pressure side inlet temperature. It is in a high-medium pressure integrated steam turbine that is heated and sent to the intermediate-pressure turbine.
[0018]
In the high-pressure turbine, the intermediate-pressure turbine, or the high-medium-pressure integrated steam turbine, at least the first stage of the body portion of the rotor shaft or the moving blade and the stationary blade is C0.05 to 0.20%, Si 0.15% or less, Mn 0.05-1.5%, Cr 9.5-13%, Ni 0.05-1.0%, V 0.05-0.35%, Nb 0.01-0.20%, N 0.005-0. From high-strength martensitic steels containing 06%, Mo 0.05-0.5%, W 1.0-4.0%, Co 2-10%, B 0.0005-0.03% and having Fe of 78% or more It is preferable to correspond to a vapor temperature of 620 to 640 ° C., or C 0.1 to 0.25%, Si 0.6% or less, Mn 1.5% or less, Cr 8.5 to 13%, Ni 0.05 to 1 0.0%, V0.05-0.5%, W0.10 0.65%, has the following Al0.1% made of a high strength martensite steel having a 80% or more of Fe, preferably corresponds to less than six hundred to six hundred twenty ° C.. The inner casing is C0.06-0.16% by weight, Si 0.5% or less, Mn 1% or less, Ni0.2-1.0%, Cr8-12%, V0.05-0.35%, Nb0. Made of high-strength martensitic steel with 85% or more Fe, including 01-0.15%, N0.01-0.8%, Mo1% or less, W1-4%, B0.0005-0.003% Is preferred.
[0019]
In the high-pressure steam turbine according to the present invention, the rotor blade has 10 stages or more, preferably the first stage is a double flow, the bearing portion of the rotor shaft has the above-mentioned requirements, and preferably the bearing center distance (L) Is 5000 mm or more (preferably 5200 to 5500 mm), the minimum diameter (D) at the portion where the stationary blade is provided is 660 mm or more (preferably 620 to 700 mm), and the (L / D) is 8.0 to 8.0 It is preferably made of high-strength martensitic steel containing 9 to 13% by weight of Cr, which is 9.0 (preferably 8.3 to 8.7).
[0020]
In the intermediate pressure steam turbine according to the present invention, the rotor blades have six or more stages symmetrically, and the bearing portion of the rotor shaft has the same requirements as described above, and preferably the first stage is implanted at the center thereof. The rotor shaft has a bearing center distance (L) of 5200 mm or more (preferably 5300 to 5800 mm) and a minimum diameter (D) at a portion where the stationary blade is provided is 620 mm or more (preferably). Is 620-680 mm), and is made of a high-strength martensitic steel containing 9-13 wt% Cr, wherein the (L / D) is 8.2-9.2 (preferably 8.5-9.0). Is preferred. In the low-pressure steam turbine having a high-pressure turbine and an intermediate-pressure turbine separately, the rotor blades have eight or more stages symmetrically, and the rotor shaft has a double flow structure in which the first stage is implanted in the center of the rotor shaft. The bearing center distance (L) is 7200 mm or more (preferably 7400 to 7600 mm) and the minimum diameter (D) at the portion where the stationary blade is provided is 1150 mm or more (preferably 1200 to 1350 mm). / D) is made of Ni-Cr-Mo-V low alloy steel containing Ni 3.25 to 4.25 wt% with 5.4 to 6.3 (preferably 5.7 to 6.1), and finally The stage blade is preferably a low-pressure steam turbine characterized by being made of high-strength martensitic steel having a value of [blade length (inch) × rotational speed (rpm)] of 125,000 or more.
[0021]
Furthermore, the present invention relates to a low pressure turbine in which a high pressure turbine and an intermediate pressure turbine are connected, and two tandem units, or a steam turbine power plant including a high intermediate pressure turbine and a low pressure turbine, the high pressure turbine and the intermediate pressure turbine, The high and medium pressure turbine has a steam inlet temperature of 600 to 660 ° C. to the first stage blade, the low pressure turbine has a steam inlet temperature of 380 to 475 ° C., and the first stage blade of the rotor shaft of the high pressure turbine is installed. And the metal temperature of the first stage rotor blade does not fall below the steam inlet temperature to the first stage rotor blade of the high pressure turbine by 40 ° C. or more (preferably 20 to 35 ° C. lower than the steam temperature). The metal temperature of the first stage rotor blade planting part and the first stage rotor blade of the rotor shaft is 75 from the steam inlet temperature to the first stage rotor blade of the intermediate pressure turbine. The rotor shaft and at least the first stage rotor blade of the high-pressure turbine, intermediate-pressure turbine, and high- and intermediate-pressure turbine contain Cr 9.5 to 13 wt%. The bearing portion of the rotor shaft is made of martensite steel and has the above-mentioned requirements. Preferably, the last stage moving blade of the low-pressure turbine has a value of [blade length (inch) × rotation speed (rpm)] of 125, It is in a steam turbine power plant characterized by being made of high-strength martensitic steel of 000 or more.
[0022]
Furthermore, the present invention provides a coal-fired boiler, a steam turbine driven by the steam obtained by the boiler, and a single machine or two or more, preferably two, that generate power of 1000 MW or more driven by the steam turbine. In the coal-fired thermal power plant equipped with a machine, the steam turbine has a high-pressure turbine, a medium-pressure turbine connected to the high-pressure turbine, and one or two low-pressure turbines, or a high-medium-pressure turbine and a low-pressure turbine. The steam inlet temperature to the first stage rotor blade is 600 to 660 ° C., and the steam inlet temperature to the first stage rotor blade is 380 to 475 ° C. 3 ° C. or more (preferably 3 to 1) from the steam inlet temperature to the first stage rotor blade of the high-pressure turbine by the boiler superheater. The steam heated to a high temperature flows into the first stage rotor blade of the high pressure turbine, and the steam exiting the high pressure turbine is moved to the first stage motion of the intermediate pressure turbine by the boiler reheater. It is heated to a temperature 2 ° C. or higher (preferably 2 to 10 ° C., more preferably 2 to 5 ° C.) higher than the steam inlet temperature to the blades and flows into the first stage blades of the intermediate pressure turbine. Steam discharged from the pressure turbine is heated to a temperature 3 ° C. or higher (preferably 3 to 10 ° C., more preferably 3 to 6 ° C.) higher than the steam inlet temperature to the first stage rotor blade of the low-pressure turbine by the boiler economizer. And the bearing portion of the rotor shaft for the high pressure turbine, intermediate pressure turbine, or high intermediate pressure turbine has the same requirements as described above, and preferably the low pressure turbine The final stage moving blade of the turbine is in the [blade length (inches) × rotational speed (rpm)] Coal-fired power plants whose value is characterized in that it consists of high-strength martensitic steel which is 125,000 or more.
[0023]
Further, in the low-pressure steam turbine described above having a high-pressure turbine and an intermediate-pressure turbine according to the present invention or having a high-medium-pressure integrated steam turbine, the steam inlet temperature to the first stage blade is 380 to 475 ° C. (preferably 400 to 450 ° C), and the rotor shaft is C0.2 to 0.3%, Si 0.05% or less, Mn 0.1% or less, Ni 3.25 to 4.25%, Cr 1.25 to 2 by weight. .25%, Mo 0.07 to 0.20%, V 0.07 to 0.2% and Fe 92.5% or more.
[0024]
In the high-pressure steam turbine described above, the rotor blade has seven or more stages (preferably 9 to 12 stages) and a blade length of 35 to 210 mm from the upstream side to the downstream side of the water vapor flow. The diameter of the implanted portion is larger than the diameter of the portion corresponding to the stationary blade, and the axial width of the implanted portion is stepwise larger in 3 steps (preferably 4 to 7 steps) on the downstream side than the upstream side, It is preferable that the ratio to the blade length is 0.6 to 1.0 (preferably 0.65 to 0.95) and decreases from the upstream side to the downstream side.
[0025]
Furthermore, in the above-described high-pressure steam turbine, the present invention has at least seven stages of the moving blades and a blade length of 35 to 210 mm from the upstream side to the downstream side of the water vapor flow, and the blade length of each adjacent stage. The ratio is 1.2 or less (preferably 1.10 to 1.15), the ratio is gradually larger on the downstream side, and the blade length is preferably larger on the downstream side than on the upstream side.
[0026]
Furthermore, in the above-described high-pressure steam turbine, the present invention has seven or more blades and a blade length of 35 to 210 mm from the upstream side to the downstream side of the water vapor flow, and corresponds to the stationary blade portion of the rotor shaft. The axial width of the portion to be reduced is smaller by two or more steps (preferably 2 to 4 steps) on the downstream side than the upstream side, and the ratio of the moving blade to the downstream blade length is 0.65 to 1.8 ( It is preferable that the ratio gradually decreases in the range of 0.7 to 1.7) toward the downstream side.
[0027]
In the above-mentioned medium pressure steam turbine, the moving blade has a double flow structure having 6 or more stages (preferably 6 to 9 stages) symmetrically and a blade part length of 100 to 300 mm from the upstream side to the downstream side of the steam flow, The diameter of the implanted portion of the rotor blade of the rotor shaft is larger than the diameter of the portion corresponding to the stationary blade, and the axial width of the implanted portion is two or more steps (preferably 3 to 6 steps) on the downstream side compared to the upstream side. ), And the ratio to the blade length is preferably 0.45 to 0.75 (preferably 0.5 to 0.7) and decreases from the upstream side to the downstream side. .
[0028]
Furthermore, the present invention is the above-described medium pressure steam turbine, wherein the moving blade has a double flow structure having six or more stages symmetrically and the blade length is 100 to 300 mm from the upstream side to the downstream side of the water vapor flow, The wing length is preferably larger on the downstream side than on the upstream side, and the ratio thereof is 1.3 or less (preferably 1.1 to 1.2) and gradually increases on the downstream side.
[0029]
Furthermore, the present invention is the above-described intermediate pressure steam turbine, wherein the rotor blade has six or more stages symmetrically left and right, and the blade length is 100 to 300 mm from the upstream side to the downstream side of the water vapor flow. The axial width of the portion corresponding to the stationary blade portion is gradually reduced in two or more steps (preferably 3 to 6 steps) on the downstream side compared to the upstream side, and the downstream blade length with respect to the downstream blade length of the moving blade It is preferable that the ratio gradually decreases as the ratio becomes lower in the range of 0.45 to 1.60 (preferably 0.5 to 1.5).
[0030]
The present invention is a low-pressure steam turbine in a power plant in which the above-described high-pressure turbine and intermediate-pressure turbine are separately provided, wherein the rotor blades are symmetrically provided with eight or more stages (preferably 8 to 10 stages), The blade portion has a length of 90 to 1300 mm from the upstream side to the downstream side of the water vapor flow, and the diameter of the implanted portion of the rotor blade of the rotor shaft is larger than the diameter of the portion corresponding to the stationary blade, and the axial direction of the implanted portion The width on the downstream side is stepwise larger than the upstream side in three steps or more (preferably 4 to 7 steps), and the ratio to the blade length is 0.15 to 1.0 (preferably 0.15). ˜0.91), it is preferable that the size decreases from the upstream side to the downstream side.
[0031]
Furthermore, the present invention provides a low-pressure steam turbine in which the above-described high-pressure turbine and intermediate-pressure turbine are separately provided, wherein the rotor blades are symmetrically provided with eight or more stages, and the blade length is from the upstream side of the steam flow. It has 90-1300 mm according to the downstream side, and the wing length of each adjacent stage is larger on the downstream side than on the upstream side, and the ratio is 1.2-1.7 (preferably 1.3-1). It is preferable that the ratio gradually increases on the downstream side in the range of .6).
[0032]
Furthermore, the present invention is the above-described low-pressure steam turbine, wherein the rotor blades are symmetrically provided with eight or more stages each having a double flow structure and the blade length is 90 to 1300 mm from the upstream side to the downstream side of the steam flow, and the rotor shaft The width in the axial direction of the portion corresponding to the stationary blade portion of the downstream blade is gradually increased in three or more steps (preferably 4 to 7 steps) on the downstream side compared to the upstream side. It is preferable that the ratio is gradually reduced as it becomes the downstream side within a range of the ratio to the part length of 0.2 to 1.4 (preferably 0.25 to 1.25).
[0033]
In the above-described high-pressure steam turbine, the moving blade has seven or more stages, and the rotor shaft has a diameter corresponding to the stationary blade smaller than a diameter corresponding to the moving blade implantation portion, and corresponds to the stationary blade. The axial width of the diameter increases stepwise in two or more stages (preferably 2 to 4 stages) on the upstream side of the water vapor flow compared to the downstream side, and the last stage of the moving blade and its front side. Is 0.75 to 0.95 times (preferably 0.8 to 0.9 times, more preferably 0.84 to 0 times) the width between the second stage and the third stage of the moving blade. 88), and the width of the rotor shaft in the axial direction of the rotor blade portion is increased in three steps or more (preferably 4 to 7 steps) on the downstream side of the water vapor flow compared to the upstream side. The axial width of the final stage of the rotor blade is larger than the axial width of the second stage. 1 to 2 times (preferably 1.4 to 1.7-fold) preferably is.
[0034]
In the above-described medium pressure steam turbine, the moving blade has six or more stages, and the rotor shaft has a diameter corresponding to the stationary blade smaller than a diameter corresponding to the moving blade implantation portion. The corresponding axial width of the diameter increases stepwise in two or more stages (preferably 3 to 6 stages) on the upstream side of the steam flow compared to the downstream side, and the final stage of the moving blade and its The width between the front and the front is 0.55 to 0.8 times (preferably 0.6 to 0.7 times) the width between the first stage and the second stage of the rotor blade, The width in the axial direction of the moving blade portion implantation portion is larger in two steps or more (preferably 3 to 6 steps) on the downstream side of the water vapor flow compared to the upstream side, and the axis of the final stage of the moving blade The width in the direction is 0.8 to 2 times (preferably 1 to 1.5 times) the axial width of the first stage. ) And it is preferably.
[0035]
In the above-described low-pressure steam turbine, the moving blade has a double flow structure with eight or more stages symmetrically, and the rotor shaft has a diameter corresponding to the stationary blade that is larger than a diameter corresponding to the moving blade implantation portion. The axial width of the diameter corresponding to the stationary blade is increased in three steps or more (preferably 4 to 7 steps) on the upstream side of the water vapor flow compared to the downstream side, and the dynamic width is increased. The width between the last stage of the blade and the front thereof is 1.5 to 2.5 times (preferably 1.7 to 2.2 times) the width between the first stage and the second stage of the moving blade. The width of the rotor shaft in the axial direction of the rotor blade implantation portion is increased stepwise in three or more stages (preferably 4 to 7 steps) on the downstream side of the water vapor flow compared to the upstream side. The axial width of the final stage of the wing is 2 to 3 times the preferred axial width of the first stage (preferably Ku is preferably 2.2 to 2.7-fold).
[0036]
The structure of the high-pressure, intermediate-pressure or high-medium-pressure integrated steam turbine and the low-pressure turbine described above can be the same structure for any of the steam temperatures of 610 to 660 ° C.
[0037]
In the rotor material of the present invention, the Cr equivalent calculated by the following formula is preferably adjusted to 4 to 8 in order to obtain high high temperature strength, low temperature toughness and high fatigue strength as a total tempered martensite structure.
[0038]
In the high-medium pressure integrated steam turbine of the present invention, the high pressure side blades have seven or more stages and the medium pressure side blades have five or more stages, and the rotor shaft has a bearing center distance (L) of 5000 mm or more (preferably 5200-5500 mm) and the minimum diameter (D) at the portion provided with the stationary blade is 660 mm or more (preferably 620-700 mm), and the (L / D) is 7.2-8.2 (preferably 7.4 to 8.0) is preferably made of high-strength martensitic steel containing 9 to 13% by weight of Cr.
[0039]
The low pressure steam turbine for the high and medium pressure integrated steam turbine of the present invention preferably has the following requirements. In the low-pressure steam turbine, the rotor blade has a five-stage or more symmetrically each, and has a double flow structure in which the first stage is implanted in the center of the rotor shaft, and the rotor shaft has a bearing center distance (L) of 5500 mm. The minimum diameter (D) at the portion (preferably 5500 to 6200 mm) and the portion provided with the stationary blade is 850 mm or more (preferably 1200 to 1350 mm), and the (L / D) is 6.3 to 7. 3 (preferably 6.5 to 7.0) Ni-Cr-Mo-V low alloy steel containing Ni 3.25 to 4.25 wt%, and the final stage blade has a blade length (inch). ) × rotational speed (rpm)] is a high strength martensitic steel having a value of 125,000 or more.
[0040]
The rotor shaft has a diameter (D) of the stationary blade portion of 850 to 1100 mm, a distance between bearing centers (L) of 6.0 to 7.0 times of D, and C0.2 to 0.3 by weight. %, Si 0.05% or less, Mn 0.1% or less, Ni 3.0 to 4.5%, Cr 1.25 to 2.25%, Mo 0.07 to 0.20%, V 0.07 to 0.2% and It consists of low alloy steel which is Fe92.5% or more.
[0041]
The moving blade has a multi-flow structure having five or more stages symmetrically and the blade length is in the range of 80 to 1300 mm from the upstream side to the downstream side of the water vapor flow, and the diameter of the implanted portion of the rotor blade on the rotor shaft is Larger than the diameter of the portion corresponding to the stationary blade, the width of the axial root portion of the implanted portion is larger than the width of the wing implanted portion at the end, and gradually increases from the downstream side to the upstream side, The ratio to the wing length is 0.25 to 1.50.
[0042]
The moving blade has a double-flow structure having five or more stages symmetrically and the blade length is in the range of 80 to 1300 mm from the upstream side to the downstream side of the water vapor flow, and the blade length of each adjacent stage is the downstream The side is larger than the upstream side, the ratio is in the range of 1.2 to 1.7, and the wing length is gradually increased on the downstream side.
[0043]
The moving blade has a multi-flow structure having 5 or more stages symmetrically and the blade length increases from the upstream side to the downstream side of the water vapor flow and is in the range of 80 to 1300 mm, and the rotor blade is implanted in the rotor shaft. The axial width of the root portion of the portion is at least three stages, and the downstream side is larger than the upstream side, and is wider than the width of the wing portion implantation portion.
[0044]
The high and medium pressure integrated steam turbine in the present invention preferably has the following configuration.
[0045]
The rotor blade on the high pressure side has seven stages or more and the blade length is 30 to 150 mm from the upstream side to the downstream side of the water vapor flow, and the diameter of the rotor blade implantation portion of the rotor shaft corresponds to the stationary blade The width of the axial root portion of the implanted portion is stepwise larger on the upstream side than on the downstream side, and the ratio to the wing length is 0.20 to 1.30 and from the upstream side to the downstream side. The intermediate pressure side blade has 5 or more stages symmetrically, the blade length is 100 to 300 mm from the upstream side to the downstream side of the water vapor flow, and the rotor blade is implanted in the rotor shaft. The diameter of the portion is larger than the diameter of the portion corresponding to the stationary blade, and the axial width of the root portion of the implanted portion is smaller than the upstream side except for the final stage, and the ratio to the length of the blade portion is 0. 40-0. It becomes smaller from the upstream side to the downstream side in 5.
[0046]
The blade has 7 or more stages and the blade length is 30 to 150 mm from the upstream side to the downstream side of the water vapor flow, and the ratio of the blade lengths of adjacent stages is 1.05 to 1.35, The blade length is gradually increased on the downstream side compared to the upstream side, the medium pressure portion has 5 or more blades, and the blade length is 100 to 1300 mm from the upstream side to the downstream side of the water vapor flow. The lengths of the adjacent blade portions are larger on the downstream side than on the upstream side, and the ratio is 1.10 to 1.30 and gradually increases on the downstream side.
[0047]
The moving blade on the high pressure side has six or more stages, and the rotor shaft has a diameter corresponding to the stationary blade smaller than a diameter corresponding to the moving blade implantation portion, The width in the axial direction is the largest at the first stage, and is gradually increased in three or more stages from the upstream side to the downstream side of the water vapor flow, the moving blade on the intermediate pressure side has five or more stages, and the rotor shaft has The diameter of the portion corresponding to the stationary blade is smaller than the diameter of the portion corresponding to the moving blade implantation portion, and the axial width of the root portion of the moving blade implantation portion is compared with the downstream side on the upstream side of the water vapor flow. The four stages are different in stages, and the first stage, the second stage, and the last stage of the moving blade are larger than the other stages.
[0048]
The turbine long blade according to the present invention has a weight ratio of C 0.08 to 0.18%, Si 0.25% or less, Mn 0.90% or less, Cr 8.0 to 13.0%, Ni 2 to 3% or less, Mo1. 5 to 3.0%, V0.05 to 0.35%, the total amount of one or two of Nb and Ta contains 0.02 to 0.20%, and N0.02 to 0.10% It is preferably made of martensitic steel.
[0049]
This steam turbine long blade must have high tensile strength and high cycle fatigue strength in order to withstand high centrifugal and vibration stresses caused by high-speed rotation. For this reason, the metal structure of the wing material must be a fully tempered martensite structure because the fatigue strength is significantly reduced if harmful δ ferrite is present.
[0050]
It is necessary for the steel of the present invention to have a component adjusted so that the Cr equivalent calculated by the formula to be described later is 10 or less so as not to substantially contain the δ ferrite phase.
[0051]
Long-wing material has a tensile strength at room temperature of 120 kgf / mm²Above, more 128.5kgf / mm²The above is preferable.
[0052]
Further, in order to obtain a homogeneous and high strength steam turbine long blade material, as a tempering heat treatment, after melting and forging, quenching is performed by holding at 1000 ° C. to 1100 ° C., preferably 0.5 to 3 hours, and then rapidly cooling to room temperature. Next, a primary tempering that is heated to 550 ° C. to 570 ° C., preferably 1-6 hours and then cooled to room temperature, and a secondary tempering that is heated to 560 ° C. to 590 ° C., preferably 1-6 hours and then cooled to room temperature Two or more tempering heat treatments are performed.
[0053]
In the present invention, the above-mentioned long blade is used as the last stage blade of the low pressure turbine, and the blade length is 914 mm (36 ") or more, and the 3600 rpm steam turbine and low pressure turbine last stage blade length is 1092 mm (43") or more. A 3000 rpm steam turbine is used, and the value of [blade length (inch) × rotation speed (rpm)] is set to 125,000 or more.
[0054]
Further, in the casing material made of the heat-resistant cast steel of the present invention, the alloy composition is adjusted so as to have a tempered martensite structure of 95% or more (δ ferrite 5% or less) to obtain high high temperature adjustment, low temperature toughness and high fatigue strength. Therefore, it is preferable to adjust the Cr equivalent calculated to 4 to 10 with the content of each element of the following formula as the weight%.
[0055]

In the 12% Cr heat resistant steel of the present invention, especially when used in steam at 621 ° C. or higher, 625 ° C., 10⁵h Creep rupture strength 10 kgf / mm²As described above, the room temperature impact absorption energy is preferably 1 kgf-m or more.
[0056]
(1) The reason for limiting the preferable component range of the 12% Cr steel used for the final stage blade of the low-pressure steam turbine in the present invention will be described.
[0057]
C needs to be at least 0.08% to obtain high tensile strength. If too much C is added, the toughness is lowered, so it must be made 0.20% or less. In particular, 0.10 to 0.18% is preferable. 0.12 to 0.16% is more preferable.
[0058]
Si is a deoxidizing agent, and Mn is a desulfurizing / deoxidizing agent that is added when steel is dissolved. Si is a δ ferrite-forming element, and if added in a large amount causes harmful δ-ferrite formation that reduces fatigue and toughness, it must be 0.25% or less. In addition, according to a carbon vacuum deoxidation method, an electroslag melting method, etc., it is not necessary to add Si, and Si addition is good. In particular, it is preferably 0.10% or less, more preferably 0.05% or less.
[0059]
A large amount of Mn reduces toughness, so it should be 0.9% or less. In particular, since Mn is effective as a deoxidizer, it is preferably 0.4% or less and more preferably 0.2% or less from the viewpoint of improving toughness.
[0060]
Cr enhances corrosion resistance and tensile strength, but if added over 13%, it causes the formation of δ ferrite structure. If less than 8%, corrosion resistance and tensile strength are insufficient, so Cr was determined to be 8-13%. From the point of strength, 10.5 to 12.5% is more preferable 11 to 12%.
[0061]
Mo has an effect of increasing tensile strength by solid solution strengthening and precipitation strengthening actions. Mo has an insufficient effect of improving the tensile strength, and if it exceeds 3%, it causes δ ferrite to be generated, so it is limited to 1.5 to 3.0%. In particular, it is preferably 1.8 to 2.7%, more preferably 2.0 to 2.5%. W and Co have the same effect as Mo.
[0062]
V and Nb precipitate carbides and increase the tensile strength, while at the same time improving the toughness. If V0.05% and Nb0.02% or less, the effect is insufficient, and if V0.35% and Nb0.2% or more, δ ferrite is generated. In particular, V is preferably 0.15 to 0.30%, more preferably 0.25 to 0.30%, and Nb is preferably 0.04 to 0.15%, more preferably 0.06 to 0.12%. Ta can be added in the same manner instead of Nb, and can be added in combination.
[0063]
Ni increases the low temperature toughness and has the effect of preventing the formation of δ ferrite. This effect is not sufficient when Ni is 2% or less, and when it exceeds 3%, the effect is saturated by addition. In particular
2.3 to 2.9% is preferable. More preferably, it is 2.4 to 2.8%.
[0064]
N is effective in improving the tensile strength and preventing the formation of δ ferrite, but if it is less than 0.02%, the effect is not sufficient, and if it exceeds 0.1%, the toughness is lowered. In particular, excellent characteristics are obtained in the range of 0.04 to 0.08, more preferably 0.06 to 0.08%.
[0065]
The reduction of Si, P and S has the effect of increasing the low temperature toughness without impairing the tensile strength, and is desirably reduced as much as possible. From the viewpoint of improving low-temperature toughness, Si is preferably 0.1% or less, P 0.015% or less, and S0.015% or less. In particular, Si of 0.05% or less, P of 0.010% or less, and S of 0.010% or less are desirable. Reduction of Sb, Sn and As also has the effect of increasing low temperature toughness, and it is desirable to reduce it as much as possible. However, from the viewpoint of the current steelmaking technology level, Sb is 0.0015% or less, Sn 0.01% or less, and As 0.02%. Limited to: In particular, Sb 0.001%, Sn 0.005%, and As 0.01% or less are desirable.
[0066]
Furthermore, in the present invention, the Mn / Ni ratio is preferably 0.11 or less.
[0067]
The heat treatment of the material of the present invention is first performed at a temperature sufficient to transform into complete austenite, uniformly heated to a minimum of 1000 ° C. and a maximum of 1100 ° C., rapidly cooled (preferably oil-cooled), and then heated to a temperature of 550 to 570 ° C. -Cooling (primary tempering), then heating and holding at a temperature of 560 to 680 ° C. to perform secondary tempering to give a fully tempered martensite structure is preferable.
[0068]
(2) Ferritic heat-resistant steel constituting the high-pressure and medium-pressure or high-medium-pressure integrated rotor shaft body, blade, nozzle, inner casing clamping bolt, and medium-pressure portion first stage diaphragm of the 620-640 ° C. steam turbine in the present invention The reason why the composition is preferably limited will be described.
[0069]
C is an element indispensable for ensuring hardenability and precipitating carbides in the tempering heat treatment process to increase the high temperature strength, and is an element necessary for 0.05% or more in order to obtain a high tensile strength. If it exceeds 0.20%, the metal structure becomes unstable when exposed to a high temperature for a long time and the creep rupture strength is lowered for a long time, so the content is limited to 0.05 to 0.20%. Desirably, it is 0.06 to 0.14%, or 0.08 to 0.13%, and 0.09 to 0.12% is particularly preferable.
[0070]
Mn is added for a deoxidizer and the like, and the effect is achieved by adding a small amount. A large amount exceeding 1.5% decreases the creep rupture strength, which is not preferable. In particular, 0.03 to 0.20% or 0.3 to 0.7% is preferable, and 0.35 to 0.65% is more preferable for the larger one. Higher strength is obtained with less Mn. In addition, workability is better when the amount of Mn is larger.
[0071]
Si is also added as a deoxidizer, but Si deoxidation is not necessary according to steelmaking techniques such as vacuum C deoxidation. Lowering Si has the effect of preventing harmful δ ferrite structure formation and preventing toughness deterioration due to segregation of grain boundaries. Therefore, when adding0.15 %Less thanIt is necessary to keep it at 0.07% or less, and particularly preferably less than 0.04%.
[0072]
Ni is a very effective element for enhancing toughness and preventing the formation of δ ferrite, but the effect is not sufficient if it is less than 0.05%, and the addition exceeding 1.0% is creep rupture strength. Is not preferable. In particular, 0.3 to 0.7%, more preferably 0.4 to 0.65% is preferable.
[0073]
Cr is an essential element for increasing high temperature strength and high temperature oxidation resistance.9%Although it is necessary, if it exceeds 13%, a harmful δ ferrite structure is formed and the high temperature strength and toughness are lowered.9~13%. In particular, 10 to 12%, more preferably 10.8 to 11.8% is preferable.
[0074]
Mo addition is performed to improve high-temperature strength. However, when it contains more than 1% W as in the steel of the present invention, addition of 0.5% or more of Mo decreases toughness and fatigue strength, so it is limited to 0.5% or less. In particular, 0.05 to 0.45%, more preferably 0.1 to 0.2% is preferable.
[0075]
W suppresses the coarsening and coarsening of carbides at high temperatures and strengthens the matrix by solid solution, so that it has the effect of remarkably increasing the high-temperature long-term strength at 620 ° C. or higher. 1 to 1.5% at 620 ° C, 1.6 to 2.0% at 630 ° C, 2.1 to 2.5% at 640 ° C, 2.6 to 3.0% at 650 ° C, 3 at 660 ° C 0.1 to 3.5% is preferable. On the other hand, if W exceeds 3.5%, δ ferrite is generated and the toughness is lowered, so it is limited to 1 to 3.5%. In particular, 2.1 to 3.0% is preferable, and 2.5 to 2.8% is more preferable.
[0076]
V has the effect of precipitating the carbonitride of V to increase the creep rupture strength, but if it is less than 0.05%, the effect is insufficient, and if it exceeds 0.3%, δ ferrite is generated and fatigue strength is increased. Reduce. In particular, 0.10 to 0.25% is preferable, and 0.15 to 0.23% is more preferable.
[0077]
Nb is a very effective element for precipitating NbC carbide and increasing the high-temperature strength. However, when added in a large amount, coarse eutectic NbC carbide is generated particularly in large steel ingots, and the strength is lowered. It is necessary to suppress it to 0.20% or less because it causes precipitation of δ ferrite which lowers the fatigue strength. If Nb is less than 0.01%, the effect is insufficient. In particular, 0.02 to 0.15%, or 0.03 to 0.10%, more preferably 0.04 to 0.08% is preferable.
[0078]
Co is an important element that distinguishes and characterizes the present invention from the conventional invention. In the present invention, the addition of Co significantly improves the high temperature strength and enhances toughness. This is considered to be due to the interaction with W and is a characteristic phenomenon in the alloy of the present invention containing 1% or more of W. In order to realize such an effect of Co, the lower limit of Co in the alloy of the present invention is 2.0%. However, even if added excessively, not only a larger effect cannot be obtained, but also the ductility is lowered. The upper limit is 10%. Desirably 2-3% for 620 ° C, 3.5-4.5% for 630 ° C, 5-6% for 640 ° C, and 6.5-7 for 650 ° C. 8% to 9% is desirable for 5% and 660 ° C.
[0079]
N is also an important element that distinguishes and characterizes the present invention from the conventional invention. N is effective in improving the creep rupture strength and preventing the formation of δ ferrite structure, but if it is 0.005% or less, the effect is not sufficient, and if it exceeds 0.05%, the toughness is reduced and the creep rupture strength is also reduced. . In particular, 0.01 to 0.03% is preferable, and 0.015 to 0.025% is more preferable.
[0080]
B is the grain boundary strength action and M₂₃C₆Solid solution in carbide, M₂₃C₆There is an effect of increasing the high temperature strength by the action of preventing the coarsening of the type carbide, and addition exceeding 0.001% is effective. However, if it exceeds 0.03%, the weldability and forgeability are impaired. Limited to ~ 0.03%. Desirably, 0.001 to 0.01% or 0.01 to 0.02% is preferable.
[0081]
Addition of Ta, Ti and Zr has an effect of increasing toughness, and a sufficient effect can be obtained by adding Ta 0.15% or less, Ti 0.1% or less and Zr 0.1% or less alone or in combination. When Ta is added in an amount of 0.1% or more, the addition of Nb can be omitted.
[0082]
In the present invention, at least the first stage of the rotor shaft body and the rotor blades and the stationary blades is C0.09 to 0.20%, Si 0.15% or less, Mn 0.05 to 1.0 with respect to the steam temperature of 620 to 630 ° C. %, Cr 9.5 to 12.5%, Ni 0.1 to 1.0%, V 0.05 to 0.30%, N 0.01 to 0.06%, Mo 0.05 to 0.5%, W2 to 3 0.5%, Co2 to 4.5%, B0.001 to 0.030%, 77% It is preferable that the steel is composed of a steel having a total tempered martensite structure having Fe. Further, it is preferable that the steel is made of steel having a total tempered martensite structure with the aforementioned Co content being 5 to 8% and 78% or more of Fe for a steam temperature of 635 to 660 ° C. In particular, high strength can be obtained by reducing the Mn content to 0.03 to 0.2% and the B content to 0.001 to 0.01% with respect to both temperatures. In particular, C 0.09 to 0.20%, Mn 0.1 to 0.7%, Ni 0.1 to 1.0%, V0.10 to 0.30%, N 0.02 to 0.05%, Mo0. Contains 0.5 to 0.5%, W2 to 3.5%, Co2 to 4% for 630 ° C. or less, Co5.5 to B0.001 to 0.01% and 630 to 660 ° C. 9.0% and B 0.01 to 0.03% are preferable.
[0083]
The Cr equivalent determined by the above formula is preferably 4 to 10.5, particularly 6.5 to 9.5 for the rotor shaft, and the other is the same.
[0084]
When the δ ferrite structure is mixed, the high pressure and medium pressure rotor materials of the steam turbine of the present invention have low fatigue strength and toughness. Therefore, the structure is preferably a uniform tempered martensite structure. In order to obtain a tempered martensite structure, the Cr equivalent calculated by the above formula must be made 10 or less by adjusting the components. If the Cr equivalent is too low, the creep rupture strength decreases, so 4 or more is preferable, and Cr equivalents of 5 to 8 are particularly preferable.
[0085]
Since the rotor body is rotated at a high speed (3000 or 3600 rpm) in steam at 621 ° C. or higher, a high stress acts on the dovetail part and the center hole part supporting the blade, so that the viewpoint of creep rupture prevention is prevented. To 10 kgf / mm²More than 10⁵h Creep rupture strength is required. In addition, since a tensile thermal stress acts on the center hole when the metal temperature is low at start-up, room temperature impact absorption energy of 1 kgf-m or more is required from the viewpoint of preventing brittle fracture.
[0086]
(3) Next, the reasons for limiting the preferred components of the rotor journal part and the low temperature region part and their functions will be described.
[0087]
C is an element necessary for obtaining a high tensile strength of 0.06% or more. However, if it exceeds 0.14%, weldability deteriorates, so 0.06 to 0.14% is preferable. In particular, 0.08 to 0.11% is preferable.
[0088]
N is effective in preventing the formation of δ ferrite structure, but the effect is not sufficient at 0.005% or less.0 . 05%Exceeding this will reduce toughness and deteriorate weldability. In particular, 0.015 to 0.03% is preferable.
[0089]
Mn is added as desulfurization, and its effect is achieved with addition of 2% or less. In particular, 1% or less is preferable.
[0090]
Si is also added as a deoxidizer, but Si deoxidation is not necessary according to steelmaking techniques such as vacuum C deoxidation. Further, by lowering Si, there is an effect of preventing generation of harmful δ ferrite structure. Therefore, when adding, it is preferable to suppress to 0.5% or less, and especially 0.2% or less is preferable.
[0091]
V has the effect of improving hardenability, but if it is 0.05% or less, the effect is insufficient, and if it exceeds 0.35%, δ ferrite may be generated and the fatigue strength may be reduced. In particular, 0.15 to 0.25% is preferable.
[0092]
Nb is an element effective for increasing toughness. However, if added in a large amount, coarse eutectic Nb carbides are formed particularly in large steel ingots. Since there is a risk of precipitation, it is preferable to keep it to 0.2% or less. Further, if Nb is 0.01% or less, the effect is insufficient. Particularly in the case of a large steel ingot, 0.03 to 0.1% is preferable, and 0.04 to 0.08% is more preferable.
[0093]
Ni is a very effective element for enhancing toughness and preventing the formation of δ ferrite, but the effect is not sufficient at 0.2% or less, and addition exceeding 2% generates a residual austenite structure. Therefore, it is not preferable.
[0094]
Cr has the effect of improving high strength and high temperature oxidation. If it exceeds 13%, a harmful δ ferrite structure will be generated, and if it is less than 7%, the oxidation resistance against high-temperature and high-pressure steam will be insufficient. Excessive Cr addition causes harmful δ ferrite structure formation and toughness reduction. In particular, 9.5 to 12%, more preferably 10.5 to 11.5% is preferable.
[0095]
W has the effect of increasing the tempering resistance and increasing the tensile strength. However, when W exceeds 3%, the toughness is lowered. 2.0 to 2.8% is preferable.
[0096]
Mo also has the effect of increasing the tempering resistance and increasing the tensile strength. Since addition of 1% or more of Mo may reduce toughness and fatigue strength, 1% or less is preferable.
[0097]
Addition of Ta, Ti and Zr has an effect of increasing toughness, and a sufficient effect can be obtained by adding Ta 0.2% or less, Ti 0.1% or less and Zr 0.1% or less alone or in combination. When Ta is added, addition of Nb can be omitted, and 0.1% or more is particularly preferable.
[0098]
Addition of B has the effect of increasing the tensile strength. If the B content exceeds 0.0030%, weldability may be significantly deteriorated, and the upper limit is preferably 0.0030%, more preferably 0.0020%. In particular, no addition is preferred.
[0099]
Co addition has the effect of preventing the precipitation of the δ ferrite phase, which is a harmful structure, and increasing the toughness. Addition exceeding 5% may reduce toughness. 1 to 3% is particularly preferable.
[0100]
In order to homogenize the entire rotor, the steel ingot weighs about 80 tons (rotor diameter: 1200 mm, length: about 8 m), so a high-level manufacturing technique is required. The rotor of the present invention melts and refines an alloy raw material with a target composition in an electric furnace, deoxidizes it by a vacuum carbon deoxidation method, produces an electrode, uses this electrode to produce a steel ingot by an electroslag remelting method, It can be produced by forming this steel ingot by hot forging. By repeating melting twice with an electric furnace and electroslag remelting method, a homogeneous rotor with little component segregation can be produced. The rotor shaft made of 12% Cr heat-resisting steel of the present invention is prepared by separately preparing two or more consumable electrodes of alloy steel with good weldability and alloy steel with high high-temperature strength as described above, and first corresponds to a journal portion. After the former consumable electrode is electroslag-melted, the latter consumable electrode corresponding to the body is immediately electroslag-melted and joined, and then the former consumable electrode is electroslag-melted and added again to obtain a predetermined predetermined length. And diameter steel ingots are obtained. Also, by preparing a consumable electrode in which the journal is made of alloy steel with good weldability (upper end and lower end) and the body (central part) is made of alloy steel with high temperature strength, this consumable electrode is electroslag-melted Can also be obtained.
[0101]
In addition, the hot forging of the steel ingot has a lower deformation resistance and is easier to forge as it is performed at a higher temperature. However, if it is performed at a too high temperature, it is cracked and must be performed in a temperature range of 850 to 1150 ° C. After hot forging and forming, it is heated to 950 to 1150 ° C., annealed, and further quenched to 1000 to 1150 ° C., and then tempered twice at 550 to 650 ° C. and 650 to 750 ° C. to give 10 kgf / mm²Above 650 ° C, 10⁵h Creep rupture strength and room temperature shock absorption energy of 1 kgf-m or more can be obtained, and a steam turbine rotor that can be used in steam of 621 ° C. or more can be produced. In addition, tempering twice can completely decompose the retained austenite and form a uniform fully tempered martensite structure.
[0102]
(4) The rotor shaft according to the present invention is formed by depositing a low-alloy steel having high bearing characteristics on the surface of the journal portion by overlay welding with a desired thickness.
[0103]
The build-up weld layer with high bearing characteristics formed in the journal portion of the steam turbine rotor shaft made of 12% Cr martensitic steel in the present invention is preferably 5 to 10 layers of a weld material made of steel. Forming a prime weld layer, and sequentially reducing the Cr amount of the weld material from the first layer to any one of the second to fourth layers, and a weld material made of steel having the same Cr amount from the fourth layer onward And the amount of Cr of the welding material used for welding of the first layer is reduced by about 2 to 6% from the amount of Cr of the base material, and the amount of Cr of the fourth and subsequent weld layers is 0.5 to 3 % (Preferably 1 to 2.5%).
[0104]
In the present invention, overlay welding is preferable for improving the bearing characteristics of the journal portion in terms of the highest safety, but the overlay welding becomes extremely difficult due to an increase in the amount of B in the steel, In order to increase the strength of the body portion by lowering the amount of B in the journal portion, the amount of B can be contained in an amount of 0.005 to 0.02%.
[0105]
The overlay weld layer obtained by the method of the present invention is preferably 5 to 10 layers. As described above, since the rapid decrease in the Cr amount as the first weld layer causes the generation of high tensile residual stress or the occurrence of weld cracking, the Cr amount as the welding material cannot be significantly reduced. Although the number of layers is increased and the Cr amount is gradually reduced, the desired Cr amount as the surface layer is further increased to 5 layers or more in order to ensure the desired thickness. Even if 10 or more layers are welded, no further effect can be obtained. As a large structural material such as a steam turbine rotor shaft, the overlay welding layer is not affected by the composition from the base material, and has a desired composition and a desired thickness. On top of that, a desired thickness is set to a desired characteristic, and the thickness is two or more layers, for example, a final finish of about 18 mm. In order to form such a thickness, even if the final finishing allowance by cutting is excluded, five overlay weld layers are preferable. It is preferable that the third and subsequent layers mainly have a tempered martensite structure and a carbide is precipitated. In particular, the composition of the weld layer after the fourth layer is C0.01 to 0.1%, Si0.3 to 1%, Mn0.3 to 1.5%, Cr0.5 to 3%, Mo0.1 by weight. Those comprising -1.5% and the balance being Fe are preferred.
[0106]
In addition, the build-up weld layer is a layer that gradually decreases the Cr content from the first layer to any of the second to fourth layers, and uses a welding rod that gradually lowers the Cr content for each layer during build-up welding. Welding, the problem of reduced ductility of the first layer weld due to the significant difference in the chromium content of the first layer weld does not occur, and it is possible to form a built-up weld layer with a desired composition without causing weld cracking. it can. As a result, the present invention can form the above-described overlay weld layer having high bearing characteristics in the final layer without extremely different chromium content between the base material and the vicinity of the first layer.
[0107]
As a welding material applied to the first layer welding, the chromium content is reduced by about 2 to 6% by weight from the chromium content of the base material. If the Cr amount of the welding material is lower than that of the base material and is 2% or less, the Cr amount of the build-up weld layer cannot be lowered sufficiently, and the effect is small. On the other hand, if it is 6% or more, it leads to a rapid decrease in Cr amount between the base metal and the overlay weld layer, and this difference in Cr amount causes a difference in thermal expansion coefficient, resulting in high tensile residual stress or occurrence of weld cracking. Cause. In addition, since the thermal expansion coefficient is smaller as the Cr content is higher, the overlay weld layer having a low Cr has a larger thermal expansion coefficient than the base material, and a high tensile residual stress is formed after welding. For this reason, welding with lower Cr steel is high in hardness due to high residual stress and causes weld cracking. Therefore, the Cr content of the welded material should be 6% or less as a value smaller than that of the base metal. is there. By using such a welding material, the chromium content of the first layer welded portion is mixed with the base material, so that it is only about 1 to 3% lower than the base material, and good welding is obtained.
[0108]
In the method of the present invention, the fourth and subsequent layers are formed using a welding material made of steel having the same Cr content. In overlay welding, up to the third layer is affected by the composition of the base metal, but since the composition of the fourth and subsequent overlay welding layers is formed only by the composition of the welding material used, the steam turbine rotor shaft A journal portion that satisfies the required characteristics can be formed. Therefore, as described above, the build-up weld layer necessary for the large structure as the steam turbine rotor shaft is about 18 mm, so that the alloy composition necessary for the final layer and the necessary and sufficient thickness for the composition are ensured. For this reason, the second and subsequent layers are welded by two or more layers with the same amount of Cr welding material, and a material satisfying the characteristics required for the above-mentioned journal portion can be formed with a sufficient thickness.
[0109]
In the present invention, the blade, nozzle, inner casing fastening bolt, and intermediate pressure first stage diaphragm are melted by vacuum melting and cast into a mold under vacuum to produce an ingot. The ingot is hot forged into a predetermined shape at the same temperature as described above, heated at 1050 to 1150 ° C. and then water-cooled or oil-quenched, then tempered at 700 to 800 ° C. Become. 10 vacuum melting^-1-10^-4Performed under mmHg. In particular, the heat-resisting steel in the present invention can be used in all stages of blades and nozzles in the high-pressure part and the medium-pressure part, but is particularly necessary in the first stage of both.
[0110]
It is preferable that the third and subsequent layers mainly have a tempered martensite structure and a carbide is precipitated. In particular, the composition of the weld layer after the fourth layer is C0.01 to 0.1%, Si 0.3 to 1%, Mn 0.3 to 1.5%, Cr 0.5 to 3%, Mo0.1 by weight. What consists of remainder Fe including -1.5% is preferable.
[0111]
(5) High pressure turbine of the present invention, medium pressure turbine and high and medium pressure turbine inner casing adjustable valve valve box, combined reheat valve valve box, main steam lead pipe, main steam inlet pipe, reheat inlet pipe, high pressure turbine nozzle box, The reason for limiting the composition of the ferritic heat-resistant steel that constitutes the first stage diaphragm of the medium pressure turbine, the high pressure turbine main steam inlet flange, the elbow, and the main steam stop valve will be explained.
[0112]
In the ferritic heat-resistant cast steel casing material, particularly by adjusting the Ni / W ratio to 0.25 to 0.75, 621 ° C., 250 kgf / cm.²625 ° C., 10 required for the above ultra-supercritical turbine high-pressure and medium-pressure inner casings and main steam stop valves and control valve casings⁵h Creep rupture strength 9 kgf / mm²As described above, a heat-resistant cast steel casing material having a room temperature impact absorption energy of 1 kgf-m or more is obtained.
[0113]
In the ferritic heat-resistant cast steel casing material of the present invention, in order to obtain high high-temperature strength, low-temperature toughness, and high fatigue strength, the Cr equivalent calculated by the above formula is preferably adjusted to 4-10.
[0114]
In the 12% Cr heat-resisting steel of the present invention, since it is used in steam at 621 ° C. or higher, 625 ° C., 10⁵h Creep rupture strength 9 kgf / mm²As described above, the room temperature impact absorption energy must be 1 kgf-m or more. Furthermore, in order to ensure higher reliability, 625 ° C., 10⁵h Creep rupture strength 10 kgf / mm²As described above, the room temperature impact absorption energy is preferably 2 kgf-m or more.
[0115]
C is an element necessary for obtaining a high tensile strength of 0.06% or more. However, if it exceeds 0.16%, the metal structure becomes unstable when exposed to high temperature for a long time, and the creep rupture strength for a long time. Is limited to 0.06 to 0.16%. In particular, 0.09 to 0.14% is preferable.
[0116]
N is effective in improving the creep rupture strength and preventing the formation of δ ferrite structure, but if it is less than 0.01%, the effect is not sufficient, and if it exceeds 0.1%, there is no significant effect. As well as the creep rupture strength. In particular, 0.02 to 0.06% is preferable.
[0117]
Mn is added as a deoxidizer, and its effect is achieved by adding a small amount. Adding a large amount exceeding 1% lowers the creep rupture strength, and 0.4 to 0.7% is particularly preferable.
[0118]
Si is also added as a deoxidizer, but Si deoxidation is not necessary according to steelmaking techniques such as vacuum C deoxidation. Further, by lowering Si, there is an effect of preventing generation of harmful δ ferrite structure. Therefore, when adding, it is necessary to suppress to 0.5% or less, and 0.1 to 0.4% is especially preferable.
[0119]
V has the effect of increasing the creep rupture strength, but if it is less than 0.05%, the effect is insufficient, and if it exceeds 0.35%, δ ferrite is generated and the fatigue strength is lowered. In particular, 0.15 to 0.25% is preferable.
[0120]
Nb is a very effective element for increasing the high-temperature strength. However, if added in a large amount, coarse eutectic Nb carbides are formed particularly in large steel ingots, and the strength is lowered or the fatigue strength is lowered. Since it causes δ ferrite to precipitate, it must be suppressed to 0.15% or less. If Nb is less than 0.01%, the effect is insufficient. Particularly in the case of a large steel ingot, 0.02 to 0.1% is preferable, and 0.04 to 0.08 is more preferable. Ni is a very effective element for enhancing toughness and preventing the formation of δ ferrite, but the effect is not sufficient if it is less than 0.2%, and the addition exceeding 1.0% is creep rupture strength. Is not preferable. 0.4 to 0.8% is particularly preferable.
[0121]
Cr has the effect of improving high strength and high temperature oxidation. If it exceeds 12%, a harmful δ ferrite structure will be formed, and if it is less than 8%, the oxidation resistance to high-temperature and high-pressure steam will be insufficient. Cr addition has the effect of increasing creep rupture strength, but excessive addition causes harmful δ ferrite structure formation and toughness reduction. In particular, it is preferably 8.0 to 10%, more preferably 8.5 to 9.5%.
[0122]
W has an effect of remarkably increasing the high temperature long time strength. When W is less than 1%, the effect is insufficient as a heat-resistant steel used at 620 to 660 ° C. On the other hand, if W exceeds 4%, the toughness is lowered. 1.0-1.5% at 620 ° C, 1.6-2.0% at 630 ° C, 2.1-2.5% at 640 ° C, 2.6-3.0% for 650 ° C At 660 ° C., 3.1 to 3.5% is preferable.
[0123]
W and Ni are correlated with each other, and by setting the Ni / W ratio to 0.25 to 0.75, a material having high strength and toughness can be obtained.
[0124]
Mo addition is performed to improve high-temperature strength. However, in the case of containing more than 1% W as in the cast steel of the present invention, addition of 1.5% or more of Mo lowers toughness and fatigue strength, so 1.5% or less is preferable, especially 0.4 to 0.8%, more preferably 0.55 to 0.70%.
[0125]
Addition of Ta, Ti and Zr has an effect of increasing toughness, and a sufficient effect can be obtained by adding Ta 0.15% or less, Ti 0.1% or less and Zr 0.1% or less alone or in combination. When Ta is added in an amount of 0.1% or more, the addition of Nb can be omitted.
[0126]
In the heat-resistant cast steel casing material of the present invention, when the δ ferrite structure is mixed, the fatigue strength and toughness are lowered. Therefore, the structure is preferably a uniform tempered martensite structure. In order to obtain a tempered martensite structure, the Cr equivalent calculated by the above formula must be made 10 or less by adjusting the components. If the Cr equivalent is too low, the creep rupture strength decreases, so it must be 4 or more. In particular, a Cr equivalent of 6 to 9 is preferable.
[0127]
Addition of B significantly increases the creep rupture strength at a high temperature (620 ° C. or higher). If the B content exceeds 0.003%, weldability deteriorates, so the upper limit is limited to 0.003%. In particular, the upper limit of the B content of the large casing is preferably 0.0028%, more preferably 0.0005 to 0.0025%, and particularly preferably 0.001 to 0.002%.
[0128]
Since the casing covers high-pressure steam at 620 ° C. or higher, high stress due to internal pressure acts. Therefore, from the viewpoint of preventing creep rupture, 10 kgf / mm²More than 10⁵h Creep rupture strength is required. Further, since thermal stress acts when the metal temperature is low at startup, room temperature impact absorption energy of 1 kgf-m or more is required from the viewpoint of preventing brittle fracture. The higher temperature side can be strengthened by containing 10% or less of Co. In particular, 1-2% for 620, 2.5-3.5% for 630 ° C, 4-5% for 640 ° C, and 5.5-6. For 5% and 660 ° C., 7 to 8% is preferable. At 600 to 620 ° C., no addition may be performed.
[0129]
In order to produce a casing with few defects, an ingot weight is as large as about 50 tons, so that advanced manufacturing technology is required. The ferritic heat-resistant cast steel casing material of the present invention can be produced by melting a raw alloy material having a target composition in an electric furnace, refining the ladle, and casting it into a sand mold. By carrying out sufficient refining and deoxidation before casting, the casting defects such as shrinkage can be reduced.
[0130]
Further, after annealing the cast steel at 1000 to 1150 ° C., tempering twice in the order of 550 to 750 ° C. and 670 to 770 ° C. in the order of normalizing heat treatment to 1000 to 1100 ° C. and quenching, 621 ° C. A steam turbine casing that can be used in the above steam can be manufactured. If the annealing and normalizing temperatures are 1000 ° C. or less, carbonitride cannot be sufficiently dissolved, and if it is too high, it causes coarsening of crystal grains. Further, the tempering twice can completely decompose the retained austenite and form a uniform tempered martensite structure. By producing by the above manufacturing method, 10 kgf / mm²Above 625 ° C, 10⁵h Creep rupture strength and room temperature shock absorption energy of 1 kgf-m or more are obtained, and a steam turbine casing usable in steam of 620 ° C. or more can be obtained.
[0131]
If O exceeds 0.015%, the high-temperature strength and toughness value are lowered, so 0.015% or less is preferable, and 0.010% or less is particularly preferable.
[0132]
The casing in the present invention has the above Cr equivalent, and the amount of δ ferrite is preferably 5% or less, more preferably 0%.
[0133]
The inner casing is preferably made of forged steel except that it is made of cast steel.
[0134]
(6) The low-pressure steam turbine rotor shaft is C0.2 to 0.3%, Si 0.1% or less, Mn 0.2% or less, Ni 3.2 to 4.0%, Cr 1.25 to 2.25% by weight. , Mo 0.1 to 0.6%, V 0.05 to 0.25% low alloy steel having a tempered bainitic structure is preferable, and is manufactured by the same manufacturing method as the above-described high pressure and medium pressure rotor shaft. preferable. In particular, the amount of Si is 0.05% or less, Mn is 0.1% or less, and other materials such as P, S, As, Sb, and Sn are used as much as possible, and the total amount is 0.025% or less. It is preferable to make a super-clean production using raw materials with few impurities. P and S are each preferably 0.010% or less, Sn, As 0.005% or less, and Sb 0.001% or less.
[0135]
(7) Other than the last stage of the low-pressure turbine blade and nozzles are C 0.05 to 0.2%, Si 0.1 to 0.5%, Mn 0.2 to 1.0%, Cr 10 to 13%, Mo 0.04 Fully tempered martensitic steel with ~ 0.2% is preferred.
[0136]
(8) Carbon cast steel having C0.2 to 0.3%, Si 0.3 to 0.7%, and Mn 1% or less is preferred for the low pressure turbine inner and outer casings.
[0137]
(9) The main steam stop valve casing and the steam control valve casing are C 0.1 to 0.2%, Si 0.1 to 0.4%, Mn 0.2 to 1.0%, Cr 8.5 to 10.5%, Mo 0.3-1.0%, W 1.0-3.0%, V 0.1-0.3%, Nb 0.03-0.1%, N 0.03-0.08%, B 0.0005-0 All tempered martensitic steel containing 0.003% is preferred.
[0138]
(10) Ti alloy is used in addition to 12% Cr steel as the final stage rotor blade of low pressure turbine, and Ti alloy having Al 5-8% and V 3-6% is used especially for length exceeding 40 inches. It is done. In particular, a high-strength material having Al 5.5 to 6.5% and V 3.5 to 4.5% for 43 inches, and Al 4 to 7%, V 4 to 7%, and Sn 1 to 3% for 46 inches is preferable.
[0139]
(11) The outer casing for high-pressure turbine, medium-pressure turbine and high-medium-pressure turbine is C0.10 to 0.20%, Si0.05 to 0.6%, Mn0.1 to 1.0%, Ni0.1 to 0 0.5%, Cr 1 to 2.5%, Mo 0.5 to 1.5%, V 0.1 to 0.35%, preferably Al 0.025% or less, B 0.0005 to 0.004% and Ti 0. It is preferable to produce the cast steel having at least one of 0.5 to 0.2% and having a total tempered bainite structure. In particular, C0.10 to 0.18%, Si0.20 to 0.60%, Mn0.20 to 0.50%, Ni0.1 to 0.5%, Cr1.0 to 1.5%, Mo0.9 A cast steel containing -1.2%, V0.2-0.3%, Al 0.001-0.005%, Ti 0.045-0.10% and B0.0005-0.0020% is preferred. More preferably, the Ti / Al ratio is 0.5-10.
[0140]
(12) C0.03-0.20% (preferably 0.03-0.15) by weight as the first stage blades of high pressure, medium pressure, high medium pressure turbines (high pressure side and medium pressure side) at a steam temperature of 625-650 ° C. %), Cr 12-20%, Mo 9-20% (preferably 12-20%), Co 12% or less (preferably 5-12%), Al 0.5-1.5%, Ti 1-3%, Fe 5% or less Ni containing 0.1% or less of Si, 0.2% or less of Mn, 0.003 to 0.015% of B, Mg 0.1% or less, rare earth element 0.5% or less, Zr 0.5% or less A base alloy can be used. The following includes 0%. After forging, it is solution treated and aged at 700-870 ° C.
[0141]
DETAILED DESCRIPTION OF THE INVENTION
Example 1
A steam temperature 600 ° C. to 649 ° C. pulverized coal direct combustion boiler and a steam turbine are required in order to improve thermal efficiency by improving the steam conditions triggered by the fuel soaring after the oil shock. An example of such a steam condition boiler is shown in Table 1.
[0142]
[Table 1]

[0143]
As the capacity increases, the pulverized coal combustion furnace becomes larger, and the furnace width is 31 m, the furnace depth is 16 m, the furnace depth is 16 m, the furnace width is 34 m, and the furnace depth is 18 m for the 1050 MW class.
[0144]
Table 2 shows the main specifications of a steam temperature of 625 ° C. and a 1050 MW steam turbine. In this example, the final stage blade length in a cross-compound type four-flow exhaust, low-pressure turbine is 43 inches, the rotational speed is 3600 rpm / min with HP-IP and 1800 rpm / min with two LP units, Is composed of the main materials shown in the table. Steam temperature of high pressure part (HP) is 625 ° C, 250kg / cm²The vapor temperature of the intermediate pressure part (IP) is heated to 625 ° C. by a reheater and is 170 to 180 kg / cm²It is operated at the pressure of The low pressure part (LP) enters at a steam temperature of 450 ° C. and is sent to the condenser with a vacuum of 100 ° C. or less and 722 mmHg.
[0145]
[Table 2]

[0146]
FIG. 1 is a cross-sectional configuration diagram of a high-pressure and intermediate-pressure steam turbine. The high-pressure steam turbine is provided with a high-pressure axle (high-pressure rotor shaft) 23 in which high-pressure rotor blades 16 are implanted in a high-pressure internal casing 18 and a high-pressure external casing 19 outside thereof. The above-mentioned high-temperature and high-pressure steam is obtained by the above-described boiler, passes through the main steam pipe, passes through the main steam inlet 28 from the flange and elbow 25 constituting the main steam inlet, and is led from the nozzle box 38 to the first stage double-flow rotor blade. It is burned. The first stage is a double flow, and eight other stages are provided on one side. A stationary blade is provided for each of these blades. The moving blade is a vertical dovetail type, double tinon, and the first stage blade length is about 35 mm. The length between the axles is about 5.25 m, the diameter of the smallest part corresponding to the stationary blade part is about 620 mm, and the ratio of the length to the diameter is about 8.5.
[0147]
The width of the rotor blade implantation part of the first stage and the last stage of the rotor shaft is almost equal, and it becomes gradually smaller in the second stage, the third stage, the fifth stage, the sixth stage, and the seventh stage to the eighth stage in accordance with the downstream side. The axial width of the second stage implant is 0.64 times that of the final stage.
[0148]
The diameter of the rotor shaft in the portion corresponding to the stator blade of the rotor shaft is smaller than that of the moving blade implantation portion. The axial width of this part is gradually reduced from the width between the second stage blade and the third stage blade to the width between the last stage blade and the preceding blade. The width of the latter is 0.86 times smaller than the width of the former. The size is reduced in two stages from the second stage to the sixth stage and from the sixth stage to the ninth stage.
[0149]
In this embodiment, the materials shown in Table 5 to be described later are made of 12% Cr steel containing no W, Co and B except that the first stage blade and nozzle are used. The length of the blade portion of the moving blade in the present embodiment is 35 to 50 mm in the first stage and becomes longer in each stage from the second stage to the last stage, and in particular, the length from the second stage to the last stage depending on the output of the steam turbine. The length is 65 to 210 mm, the number of stages is 9 to 12, and the length of the wing portion of each stage is longer at a rate of 1.10 to 1.15 with the downstream side being adjacent to the upstream side. The ratio gradually increases on the downstream side.
[0150]
The medium-pressure steam turbine rotates the generator together with the high-pressure steam turbine by the steam heated from the high-pressure steam turbine to 625 ° C. again by the reheater, and is rotated at a rotational speed of 3600 times / min. . Like the high-pressure turbine, the intermediate pressure turbine has an intermediate pressure inner casing 21 and an outer casing 22, and a stationary blade is provided in opposition to the intermediate pressure rotor blade 17. The rotor blades 17 have six stages and two flows, and are provided on the left and right sides almost symmetrically with respect to the longitudinal direction of the medium pressure axle (medium pressure rotor shaft). The distance between the bearing centers is about 5.5 m, the first stage blade length is about 92 mm, and the last stage blade length is about 235 mm. Dovetil is a reverse chestnut type. The diameter of the rotor shaft corresponding to the stationary blade before the final stage moving blade is about 630 mm, and the ratio of the distance between the bearings to the diameter is about 8.7 times.
[0151]
The rotor shaft of the medium pressure steam turbine of this embodiment has a stepwise increase in the axial width of the rotor blade implantation portion in three stages from the first stage to the fourth stage, the fifth stage, and the last stage. It is about 1.4 times larger than that.
[0152]
Further, the diameter of the portion of the rotor shaft of the steam turbine corresponding to the stationary blade portion is small, and the width thereof is gradually reduced in four stages according to the first stage blade, the second stage, and the last stage blade. Therefore, the width in the latter axial direction relative to the former is reduced to about 0.7 times.
[0153]
In this embodiment, 12% Cr steel not containing W, Co and B is used except that the materials shown in Table 5 described later are used for the first stage blade and nozzle. The length of the blade portion of the moving blade in this embodiment becomes longer in each stage from the first stage to the last stage, and the length from the first stage to the last stage is 90 to 350 mm depending on the output of the steam turbine. In the stage, the length of the wing part of each stage is longer at a ratio of 1.1 to 1.2, with the downstream side being adjacent to the upstream side.
[0154]
The diameter of the implanted portion of the moving blade is larger than that of the portion corresponding to the stationary blade, and the width of the implanted portion increases as the blade length of the moving blade increases. The ratio of the width to the blade length of the moving blade is 0.5 to 0.7 from the first stage to the last stage, and gradually decreases from the first stage to the last stage.
[0155]
FIG. 2 is a sectional view of the low-pressure turbine. The low-pressure turbine is coupled to two tandems and has almost the same structure. Each of the moving blades 41 has eight stages on the left and right sides and is substantially symmetric with respect to the left and right. The last blade length is 43 inches. No. 7 12% Cr steel is used, and has the double tinon and saddle type dovetail shown in FIG. 3, and the nozzle box 44 is a double flow type. The rotor shaft 43 is a tempered bainite of a super clean material made of Ni 3.75%, Cr 1.75%, Mo 0.4%, V 0.15%, C 0.25%, Si 0.05%, Mn 0.10%, residual Fe. Forged steel with a texture is used. For the moving blades and stationary blades other than the final stage, 12% Cr steel containing 0.1% Mo is used. C0.25% cast steel is used for the inner and outer casing materials. The distance between the centers of the bearings 43 in this embodiment is 7500 mm, the diameter of the rotor shaft corresponding to the stationary blade portion is about 1280 mm, and the diameter at the moving blade implantation portion is 2275 mm. The distance between the bearing centers for this rotor shaft diameter is about 5.9.
[0156]
3 is a perspective view of a long blade of 1092 mm (43 ″). 51 is a blade portion where high speed steam strikes, 52 is a blade implantation portion on the rotor shaft, and 53 is a pin for supporting the centrifugal force of the blade. Holes 54 are erosion shields (Co-base alloy stellite plates are joined by welding) to prevent erosion due to water droplets in the steam. The 43 ″ long blades in this example were melted by electroslag remelting. Then, forging heat and treatment were performed.
[0157]
Table 3 shows the chemical composition (% by weight) of 12% Cr steel related to the steam turbine long blade material. Each sample was melted by electroslag after 150 kg vacuum arc, heated to 11150 ° C. and forged as an experimental material. Sample No. 1 was heated to 1000 ° C. for 1 h, then cooled to room temperature by oil quenching, then heated to 570 ° C. and held for 2 h, then air cooled to room temperature. No. 2 was heated to 1050 ° C. for 1 h, cooled to room temperature by oil quenching, then heated to 570 ° C., held for 2 h, and then air-cooled to room temperature. Sample No. 3-No. 6 is cooled to room temperature by oil quenching after heating at 1050 ° C. for 1 h, then heated to 560 ° C. and held for 2 h, then air cooled to room temperature (primary tempering), further heated to 580 ° C. and held for 2 h, then cooled to the room temperature (Secondary tempering).
[0158]
In Table 3, no. 3, 4 and 5 are materials of the present invention, No. No. 6 is a comparative material and no. 1 and 2 are current 26 ″ long wing materials.
[0159]
Table 4 shows the room temperature mechanical properties of these samples. This invention material (No. 3-5) is the tensile strength (120 kgf / mm) requested | required as a long blade material for steam turbines.²Above or 128.5kgf / mm²Above) and low temperature toughness (20 ° C. V-notch Charpy impact value 2.5 kgf-m / cm)²It was confirmed that the above was sufficiently satisfied.
[0160]
On the other hand, the comparative material No. When 1 and 6 are used for steam turbine long blades, the values indicated by the tensile strength and impact value are low. Comparative material trial No. 2 has low tensile strength and toughness. No. 5, the impact value is 3.8 kgf-m / cm²4kgf-m / cm for 43 "and above²The above requirements are slightly lacking.
[0161]
[Table 3]

[0162]
[Table 4]

[0163]
In this example, the Ni and Mo contents are included at the same content to increase both the strength and toughness at low temperatures, and the strength tends to decrease as the difference between the two contents increases. . When the Ni amount is 0.6% or more less than the Mo amount, the strength rapidly decreases. Conversely, when the Ni amount is 1.0% or more, the strength rapidly decreases. Therefore, (Ni-Mo) amount of -0.6 to 1.0% indicates high strength. The (Ni—Mo) amount shows a high value before and after the impact value decreases near −0.5%.
[0164]
The quenching temperature is 975 to 1125 ° C and the primary tempering temperature is 550 to 560 ° C, and the secondary tempering temperature is 560 to 590 ° C. Characteristics required for long blade materials (tensile strength ≧ 128.5 kgf / mm², 20 ° C. V-notch Charpy impact value ≧ 4 kgf-m / cm²).
[0165]
The 12% Cr steel in this example has a tensile strength of 120 kgf / mm as described above.²Above and impact value 4kgf-m / cm²Although what has the above is preferable, what makes an impact value (y) more than the value calculated | required by [-0.45x (tensile strength) +61.5] is especially preferable.
[0166]
In particular, the 12% Cr steel according to the present invention has a C + Nb amount of 0.18 to 0.35%, an (Nb / C) ratio of 0.45 to 1.00, and an (Nb / N) ratio of 0.8. ~ 3.0 is preferred.
[0167]
In the low-pressure turbine of the present embodiment, the axial width of the moving blade implantation portion is gradually increased in four stages from the first stage to the third stage, the fourth stage, the fifth stage, the sixth stage to the seventh stage, and the eighth stage. Is about 2.5 times larger than the width of the first stage.
[0168]
In addition, the diameter of the part corresponding to the stationary blade part is small, and the axial width of that part is gradually increased from the first stage moving blade side to the fifth, sixth and seventh stages. The width on the final stage side is about 1.9 times larger than that on the first stage side.
[0169]
The moving blades in this embodiment have 8 stages, and the length of the blades is 3 "to 5", 7 ", 10", 13 ", 18", 27 "and 43" in the first stage. Depending on the output of the steam turbine, the length from the first stage to the last stage is 90 to 1270 mm, it is 8 stages or 9 stages, the blade length of each stage is the length adjacent to the upstream side on the downstream side It is long at a rate of 1.3 to 1.6 times.
[0170]
The diameter of the implanted portion of the moving blade is larger than that of the portion corresponding to the stationary blade, and the width of the implanted portion increases as the blade length of the moving blade increases. The ratio of the width to the blade length of the moving blade is 0.15 to 0.91 from the first stage to the last stage, and is gradually reduced from the first stage to the last stage.
[0171]
Further, the width of the rotor shaft at the portion corresponding to each stationary blade is gradually reduced at each stage from between the first stage and the second stage to between the last stage and the front thereof. The ratio of the width to the blade length of the moving blade is 0.25 to 1.25, and decreases from the upstream side to the downstream side.
[0172]
In addition to this embodiment, the same configuration is applied to a 1000 MW class large-capacity power plant in which the steam inlet temperature to the high-pressure steam turbine and the medium-pressure steam turbine is 610 ° C. and the steam inlet temperature to the two low-pressure steam turbines is 385 ° C. It can be.
[0173]
FIG. 4 shows a typical plant configuration diagram of a coal combustion high-temperature high-pressure steam turbine plant.
[0174]
The high-temperature high-pressure steam turbine plant in this embodiment is mainly composed of a coal-fired boiler 51, a high-pressure turbine 52, an intermediate-pressure turbine 53, a low-pressure turbine 54, a low-pressure turbine 55, a condenser 56, a condensate pump 57, a low-pressure feed water heater system 58, It comprises a deaerator 59, a booster pump 60, a feed water pump 61, a high pressure feed water heater system 63, and the like. That is, the super high temperature and high pressure steam generated in the boiler 51 enters the high pressure turbine 52 to generate power, and then is reheated again in the boiler 51 and enters the intermediate pressure turbine 53 to generate power. The intermediate-pressure turbine exhaust steam enters the low-pressure turbines 54 and 55 to generate power, and then condenses in the condenser 56. This condensate is sent to a low-pressure feed water heater system 58 and a deaerator 59 by a condensate pump 57. The feed water deaerated by the deaerator 59 is sent to the high-pressure feed water heater 63 by the booster pump 60 and the feed water pump 61 to be heated, and then returns to the boiler 51.
[0175]
Here, in the boiler 51, the feed water passes through the economizer 64, the evaporator 65, and the superheater 66 to become high-temperature and high-pressure steam. On the other hand, the boiler combustion gas that has heated the steam exits the economizer 64 and then enters the air heater 67 to heat the air. Here, the feed water pump 61 is driven by a feed water pump drive turbine that is operated by extracted steam from a medium pressure turbine.
[0176]
In the high-temperature high-pressure steam turbine plant configured in this way, the temperature of the feed water leaving the high-pressure feed water heater system 63 is much higher than the feed water temperature in the conventional thermal power plant. The temperature of the combustion gas exiting the economizer 64 is much higher than that of the conventional boiler. For this reason, heat recovery from this boiler exhaust gas is carried out so as not to lower the gas temperature.
[0177]
The same high-pressure turbine, medium-pressure turbine, and one or two low-pressure turbines may be connected to the tandem instead of the present embodiment, and the same can be applied to a tandem compound power plant that generates power by rotating one generator. Can be configured. As in this embodiment, a generator shaft having an output of 1050 MW has a higher strength as the generator shaft. In particular, C0.15-0.30%, Si0.1-0.3%, Mn0.5% or less, Ni3.25-4.5%, Cr2.05-3.0%, Mo0.25-5. It has a total tempered bainite structure containing 60%, V0.05-0.20%, and room temperature tensile strength 93 kgf / mm²Above, especially 100kgf / mm²As described above, 50% FATT is preferably 0 ° C. or less, particularly −20 ° C. or less, and the magnetization force at 21.2 KG is 985 AT / cm or less, and the total amount of P, S, Sn, Sb, As as impurities Is preferably 0.025% or less and the Ni / Cr ratio is 2.0 or less.
[0178]
FIG. 5 is a front view of a high pressure and FIG. 6 is a front view of an intermediate pressure turbine rotor shaft. The high-pressure turbine shaft shown in FIG. 5 has a structure in which eight stages of blades are planted around the first stage blade planting portion on the multi-stage side. The medium pressure turbine shaft has multistage blades provided with blade planting portions approximately symmetrically on the left and right sides, each having a substantially centered border. The rotor shaft for the low-pressure turbine is not shown in the figure, but any rotor shaft of high-pressure, medium-pressure, or low-pressure turbine has a center hole, and through this center hole there is no defect by ultrasonic inspection, visual inspection and fluorescence flaw detection Is inspected.
[0179]
Table 5 shows chemical compositions (% by weight) used in the main parts of the high-pressure turbine, intermediate-pressure turbine, and low-pressure turbine of this example. In this example, the high-temperature part and the high-pressure part of the medium-pressure part all have a thermal expansion coefficient of 12 × 10 having a ferrite crystal structure.^-6Since there was no difference in thermal expansion coefficient, there was no problem at all.
[0180]
The rotor of the high-pressure part and the medium-pressure part is prepared by melting 30 tons of heat-resistant cast steel according to the body part and bearing part shown in Table 5 in an electric furnace, decarbonizing it in a vacuum, casting it in a mold, and forging it. A rod was prepared, and using this electrode rod, the bearing portion was first electroslag-melted, and immediately after that, the body portion was electro-remelted. 3700 mm) and forged. This forging was performed at a temperature of 1150 ° C. or lower in order to prevent forging cracks. Further, this forged steel was heated to 1050 ° C. after annealing and heat-treated by water spray cooling and quenching twice at 570 ° C. and 690 ° C., and obtained by cutting into the shapes shown in FIGS. The body portion and the bearing portion are joined at the position indicated by the dotted line. As shown in FIG. 5, in the rotor shaft for the high pressure steam turbine, between the downstream final stage of the blade and the front thereof, and in the rotor shaft for the intermediate pressure steam turbine shown in FIG. 6, between the downstream final stage and the front thereof, respectively. It is joined. In the present embodiment, the upper side of the electroslag steel ingot is set to the first stage blade side of the trunk and the lower side is set to the final stage side. The blades and nozzles of the high-pressure part and medium-pressure part are similarly forged into a blade and nozzle material shape (width 150 mm, height 50 mm, length 1000 mm) by melting the heat-resistant steel listed in Table 5 in a vacuum arc melting furnace. And molded. This forging was performed at a temperature of 1150 ° C. or lower in order to prevent forging cracks. The forged steel is heated to 1050 ° C., subjected to oil quenching, tempered at 690 ° C., and then cut into a predetermined shape.
[0181]
The inner casing, main steam stop valve casing, and steam control valve casing of the high-pressure part and the intermediate-pressure part were prepared by melting the heat-resistant cast steel shown in Table 5 in an electric furnace, refining the ladle, and casting it into a sand mold. By carrying out sufficient refining and deoxidation before casting, a product having no casting defects such as shrinkage cavities was obtained. The weldability evaluation using this casing material was performed according to JIS Z3158. Preheating, between passes, and after heat start temperature were 200 ° C., and post heat treatment was 400 ° C. × 30 minutes. The material of the present invention had no weld cracks and good weldability.
[0182]
[Table 5]

[0183]
Table 6 shows the mechanical properties and heat treatment conditions obtained by cutting the above-mentioned ferritic steel high-temperature steam turbine main members.
[0184]
As a result of investigating the center portion of the rotor shaft, the characteristics required for the high-pressure and medium-pressure turbine rotor (625 ° C., 10⁵h Strength ≧ 13kgf / mm², 20 ° C. impact absorption energy ≧ 1.5 kgf−m) was sufficiently satisfied. This demonstrated that a steam turbine rotor that can be used in steam at 620 ° C. or higher can be manufactured. Further, as a result of investigating the characteristics of this blade, the characteristics required for the first stage blade of a high-pressure / medium-pressure turbine (625 ° C., 10⁵h Strength ≧ 15kgf / mm²) Was sufficiently satisfied. This demonstrated that a steam turbine blade that can be used in steam at 620 ° C. or higher can be produced.
[0185]
Furthermore, as a result of investigating the characteristics of this casing, the characteristics required for high-pressure and medium-pressure turbine casings (625 ° C., 10⁵h Strength ≧ 10kgf / mm², 20 ° C. impact absorption energy ≧ 1 kgf−m) was sufficiently satisfied and welding was confirmed. This demonstrated that a steam turbine casing that can be used in steam at 620 ° C. or higher can be produced.
[0186]
[Table 6]

[0187]
In this example, Cr—Mo low alloy steel was overlay welded to the journal portion of the rotor shaft to improve bearing characteristics. Overlay welding is as follows.
[0188]
As a test welding rod, a coated arc welding rod (diameter 4.0φ) was used. Table 7 shows the chemical composition (% by weight) of the deposited metal that was welded using the welding rod. The composition of the deposited metal is almost the same as the composition of the welding material.
[0189]
The welding conditions are a welding current of 170 A, a voltage of 24 V, and a speed of 26 cm / min.
[0190]
[Table 7]

[0191]
As shown in Table 8 on the surface of the above-described specimen base material, the overlay welding was performed as No.1. 1 and no. 8 layers were welded by combining the welding rods used for each layer for the two types. The thickness of each layer was 3-4 mm, the total thickness was about 28 mm, and the surface was ground about 5 mm.
[0192]
The welding conditions are preheating, between passes, stress relief annealing (SR) start temperature is 250 to 350 ° C., and SR treatment condition is 630 ° C. × 36 hours.
[0193]
[Table 8]

[0194]
In order to confirm the performance of the welded portion, overlay welding was similarly performed on the plate material, and a 160 ° side bending test was performed, but no cracks were observed in the welded portion. In any case, the sixth and subsequent layers have the compositions shown in the respective tables.
[0195]
Furthermore, a bearing sliding test by rotation according to the present invention was performed, and none of the bearings had an adverse effect on the bearing and was excellent in oxidation resistance.
[0196]
In place of the present embodiment, a high pressure steam turbine, an intermediate pressure steam turbine, and two low pressure steam turbines are connected to a tandem, and a tandem type power plant having 3600 revolutions can be similarly configured.
[0197]
(Example 2)
Table 9 shows the main specifications of the 600MW steam turbine with a steam temperature of 621 ° C. This example is a tandem compound double flow type, the last stage blade length in a low pressure turbine is 43 inches, HP / IP integrated type and one LP unit have a rotation speed of 3000 rpm / min, and shown in the table in the high temperature part Consists of main materials. Steam temperature of the high pressure part (HP) is 600 ° C, 250 kg / cm²The steam temperature of the intermediate pressure part (IP) is heated to 600 ° C. by a reheater and is 170 to 180 kg / cm²It is operated at the pressure of The low pressure part (LP) enters at a steam temperature of 450 ° C. and is sent to the condenser with a vacuum of 100 ° C. or less and 722 mmHg.
[0198]
[Table 9]

[0199]
FIG. 7 is a cross-sectional view of a high-pressure / medium-pressure integrated steam turbine, and FIG. 8 is a cross-sectional view of the rotor shaft. The high pressure side steam turbine is provided with a high and medium pressure axle (high pressure rotor shaft) 23 in which a high pressure side moving blade 16 is implanted in an inner casing 18 and an outer casing 19 outside thereof. The above-mentioned high-temperature and high-pressure steam is obtained by the above-described boiler, passes through the main steam pipe, passes through the main steam inlet 28 from the flange and elbow 25 constituting the main steam inlet, and is led from the nozzle box 38 to the first stage blade. . The rotor blades are provided in 8 stages on the high pressure side on the left side in the figure and 6 stages on the intermediate pressure side (about half of the right side in the figure). A stationary blade is provided for each of these blades. The moving blade is a vertical dovetil type, double tinon, high pressure side first stage blade length of about 40 mm, and medium pressure side first stage blade length is 130 mm. The length between the bearings 43 is about 5.7 m, the diameter of the smallest part corresponding to the stationary blade part is about 740 mm, and the ratio of the length to the diameter is about 7.7. A central hole is also provided in the high and medium pressure integrated rotor shaft, and the presence or absence of defects is inspected.
[0200]
The first stage and the last stage of the rotor blade implantation root part of the high pressure side rotor shaft are the widest in the first stage, the second stage to the seventh stage are smaller than that, 0.40 to 0.56 times the first stage, and both are equivalent. The final stage is between the first stage and the second to seventh stages, and is 0.46 to 0.62 times larger than the first stage.
[0201]
On the high pressure side, the blade and nozzle are made of 12% Cr steel shown in Table 5 to be described later. The length of the blade portion of the moving blade in the present embodiment is 35 to 50 mm in the first stage and becomes longer in each stage from the second stage to the last stage, and in particular, the length from the second stage to the last stage depending on the output of the steam turbine. Is in the range of 50 to 150 mm, the number of stages is in the range of 7 to 12 stages, and the length of the wing portion of each stage is 1.05 to 1. While it becomes long within the range of 35 times, the ratio is gradually increased on the downstream side.
[0202]
The medium-pressure steam turbine rotates the generator together with the high-pressure steam turbine by steam heated again by the reheater to 600 ° C. The steam discharged from the high-pressure steam turbine is rotated at a rotational speed of 3000 rpm. The intermediate pressure side turbine has an internal casing 21 and an external casing 22 as in the high pressure side turbine, and a stationary blade is provided in opposition to the moving blade 17. The moving blade 17 has six stages. The first stage blade length is about 130 mm, and the last stage blade length is about 260 mm. Dovetil is a reverse chestnut type. The diameter of the rotor shaft corresponding to the stationary blade is about 740 mm.
[0203]
The rotor shaft of the high and medium pressure steam turbine has the largest axial width of the root of the rotor blade implantation at the first stage, the second stage is smaller than that, the 3rd to 5th stages are smaller than the second stage, and both are the same. The width is the size between the third and fifth stages and the second stage, and is 0.48 to 0.64 times the first stage. The first stage is 1.1 to 1.5 times the second stage.
[0204]
On the medium pressure side, a 12% Cr steel shown in Table 5 described later is used for the blade and nozzle. The length of the blade portion of the moving blade in this embodiment becomes longer in each stage from the first stage to the last stage, and the length from the first stage to the last stage is 90 to 350 mm and the number of stages is 6 to 6 depending on the output of the steam turbine. Within the range of 9 stages, the length of the wings of each stage is longer at a rate of 1.10 to 1.25, with the downstream side being adjacent to the upstream side. The implanted portion of the moving blade has a larger diameter than the portion corresponding to the stationary blade, and its width is related to the length and position of the moving blade. The ratio of the width to the blade length of the moving blade is the largest at the first stage, 1.35 to 1.80 times, the second stage is 0.88 to 1.18 times, and the third to sixth stages are the final stage. It is small, 0.40 to 0.65 times. FIG. 9 is a sectional view of the low-pressure turbine, and FIG. 10 is a sectional view of the rotor shaft. One low-pressure turbine is tandemly coupled to high and medium pressure. The moving blade 41 has six stages on the left and right sides, is substantially symmetrical, and a stationary blade 42 is provided corresponding to the moving blade. The final stage blade length is 43 inches, 12% Cr steel or Ti-base alloy shown in Table 3 is used, and all of them shown in FIG. 3 have double tinon and vertical dovetail. It is a double flow type. Ti-based alloys are age-hardened and contain 6% Al and 4% V by weight. Rotor shaft 43 is a total tempered bainite of super clean material consisting of Ni 3.75%, Cr 1.75%, Mo 0.4%, V 0.15%, C 0.25%, Si 0.05%, Mn 0.10%, residual Fe. Forged steel with a texture is used. For the moving blades and stationary blades other than the final stage, 12% Cr steel containing 0.1% Mo is used. C0.25% cast steel is used for the inner and outer casing materials. The distance between the centers of the bearings 43 in this embodiment is 6000 mm, the diameter of the rotor shaft corresponding to the stationary blade portion is about 890 mm, and the diameter at the moving blade implantation portion is 1820 mm at the same stage. The distance between the bearing centers with respect to the rotor shaft diameter corresponding to the stator blade is about 6.7.
[0205]
In the low-pressure turbine, the axial width of the root portion of the rotor blade implantation is the smallest at the first stage, the second and third stages are the same on the downstream side, the fourth and fifth stages are the same, and gradually increase in four stages. The width is 3.8 to 4.8 times larger than the width of the first stage. The second and third stages are 1.05 to 1.40 times the first stage, the fourth and fifth stages are 1.05 to 1.35 times the second and third stages, and the last stage is 2.8 to 3.2, which are fourth and fifth stages. It has doubled. The width of the root portion is indicated by a point connecting a diverging extension line and the diameter of the rotor shaft.
[0206]
The blade length of the moving blade in the present embodiment becomes longer at each stage as it reaches the final stage of 5 "to 7", 9 ", 16", 26 "and 43" of the first stage, and the first stage depends on the output of the steam turbine. The length of the last stage is within a range of 80 to 1270 mm, with a maximum of 9 stages, and the wing length of each stage is 1.3 to 1.9 times that the downstream side is adjacent to the upstream side It is getting longer inside.
[0207]
The root portion of the moving blade has a larger diameter than the portion corresponding to the stationary blade, and its width increases as the blade length of the moving blade increases. The ratio of the width to the blade length of the moving blade is 0.30 to 1.5 excluding the first stage to the last stage, and gradually decreases from the first stage to the last stage. It is smaller in the range of 0.15 to 0.40. The final stage has a ratio of 0.55 to 0.65.
[0208]
The final stage moving blade in the present embodiment is the same as that in the first embodiment.
[0209]
In addition to this embodiment, the same applies to a 1000 MW class large capacity power plant in which the steam inlet temperature of the high and medium pressure steam turbine is 610 ° C or higher, the steam inlet temperature to the low pressure steam turbine is about 450 ° C, and the outlet temperature is about 60 ° C. It can be configured.
[0210]
The high-temperature high-pressure steam turbine power plant in this embodiment is mainly a boiler, a high-medium pressure turbine, a low-pressure turbine, a condenser, a condensate pump, a low-pressure feed water heater system, a deaerator, a boost pump, a feed pump, and a high-pressure feed water heater system. Etc. That is, the super high temperature and high pressure steam generated in the boiler enters the high pressure turbine and generates power, and then is reheated again in the boiler to enter the intermediate pressure turbine and generate power. The high and medium pressure turbine exhaust steam enters the low pressure turbine to generate power, and then condenses in the condenser. This condensate is sent to a low-pressure feed water heater system and a deaerator by a condensate pump. The feed water deaerated by the deaerator is sent to a high-pressure feed water heater by a booster pump and a feed water pump, and after being heated, returns to the boiler.
[0211]
Here, in the boiler, the feed water passes through a economizer, an evaporator, and a superheater and becomes high-temperature and high-pressure steam. On the other hand, the boiler combustion gas that has heated the steam exits the economizer and enters the air heater to heat the air. Here, a feed water pump driving turbine that is operated by extracted steam from an intermediate pressure turbine is used to drive the feed water pump.
[0212]
In the high-temperature and high-pressure steam turbine plant configured in this way, the temperature of the feed water leaving the high-pressure feed water heater system 63 is much higher than the feed water temperature in the conventional thermal power plant. The temperature of the combustion gas exiting the economizer is much higher than that of the conventional boiler. For this reason, heat recovery from this boiler exhaust gas is carried out so as not to lower the gas temperature.
[0213]
In this embodiment, a tandem compound double flow power plant is constructed in which a high-medium pressure turbine and one low-pressure turbine are connected to one generator tandem to generate power. As another example, a generator shaft having an output of 1050 MW has a higher strength as the generator shaft. In particular, C0.15-0.30%, Si0.1-0.3%, Mn0.5% or less, Ni3.25-4.5%, Cr2.05-3.0%, Mo0.25-5. It has a total tempered bainitic structure containing 60%, V0.05-0.20%, and a room temperature tensile strength of 93 kgf / mm²Above, especially 100kgf / mm²As described above, 50% FATT is preferably 0 ° C. or less, particularly −20 ° C. or less, and the magnetization force at 21.2 KG is 985 AT / cm or less, and the total amount of P, S, Sn, Sb, As as impurities Is preferably 0.025% or less and the Ni / Cr ratio is 2.0 or less.
[0214]
Table 5 above shows the chemical composition (% by weight) used in the main part of the high and medium pressure turbine and the low pressure turbine of this example. In this example, the high temperature part in which the high-pressure side and the medium-pressure side are integrated, No. 4 of Example 4 described later. A thermal expansion coefficient of 12 × 10 having a ferrite-based crystal structure is used except that the martensitic steel of No. 9 is used and the steel of Table 5 is used.^-6Since there was no difference in thermal expansion coefficient, there was no problem at all.
[0215]
The rotor shaft body for the high and medium pressure steam turbine uses a 12% Cr heat resistant steel shown in Table 10, and in this example, No. Using the heat-resistant cast steel described in No. 2, the production of the body part and the bearing part electrode as in Example 1 and the production of a steel ingot corresponding to the rotor shaft from the electrode, and the rotor shape (diameter 1450 mm, length 5000 mm) ) And forged. This forging was performed at a temperature of 1150 ° C. or lower in order to prevent forging cracks. Further, this forged steel was heated to 1050 ° C. after annealing and heat-treated by water spray cooling and quenching twice at 570 ° C. and 690 ° C., and obtained by cutting into the shape shown in FIG. The body part and the bearing part were manufactured with different materials at the dotted line part shown in the figure. Furthermore, overlay welding to the bearing journal portion 45 was performed in the same manner.
[0216]
(Example 3)
An alloy having the composition shown in Table 10 was cast into a 10 kg ingot by vacuum melting and forged to 30 mm square. In the case of a large steam turbine rotor shaft, 1050 ° C. × 5 hours 100 ° C./h cooling quenching, 570 ° C. × 20 hours primary tempering, 690 ° C. × 20 hours secondary tempering, The blade was quenched at 1100 ° C. × 1 hour and tempered at 750 ° C. × 1 hour, 625 ° C., 30 kgf / mm²A creep rupture test was carried out. The results are shown in Table 7.
[0217]
No. in Table 10 1-No. It can be seen that the present invention alloy No. 6 is preferable for application to steam conditions of 620 ° C. or higher and has a long creep rupture life. The creep rupture time improves as the amount of Co increases. However, a large amount of Co tends to cause heat embrittlement when heated at 600 to 660 ° C. Therefore, 620 to 630 is required to increase both strengthening and toughness. It is preferably 2 to 5% with respect to ° C and 5.5 to 8% with respect to 630 to 660 ° C. B is 0.03% or less, indicating an excellent strength. At 620 to 630 ° C., the B amount is 0.001 to 0.01% and the Co amount is 2 to 4%. On the higher temperature side of 630 to 660 ° C., the B amount is 0.01 to 0.03%, and the Co amount is By increasing the content to 5 to 7.5%, high strength can be obtained.
[0218]
It has been clarified that N is less strengthened at temperatures exceeding 600 ° C. in this example, and the strength is higher than that of N having a large amount. N amount is preferably 0.01 to 0.04%. In vacuum melting, N is hardly contained, so it is added by a master alloy.
[0219]
As shown in Table 10, the rotor material is No. in this example. Corresponds to the alloy No. 2 and high strength is obtained. If the Mn content is reduced to 0.09%, the same Co content is obtained, which indicates a higher strength. Therefore, for further strengthening, the Mn content is preferably 0.03 to 0.20%.
[0220]
[Table 10]

[0221]
Example 4
Table 11 shows the chemical composition (% by weight) of the inner casing material for high pressure, medium pressure and high / medium pressure turbines of the present invention. The sample was assumed to be a thick part of a large casing, 200 kg was melted using a high-frequency induction melting furnace, and cast into a sand mold having a maximum thickness of 200 mm, a width of 380 mm, and a height of 440 mm to produce an ingot. Samples were annealed at 1050 ° C. × 8 h for furnace cooling, and then normalized (1050 ° C. × 8 h → air cooling) and tempered (710 ° C. × 7 h → air cooling, 710 ° C. × 7 h) (→ Air-cooled twice).
[0222]
Weldability evaluation was performed according to JIS Z3158. Preheating, between passes, and after heat start temperature were set to 150 ° C., and post heat treatment was set to 400 ° C. × 30 minutes.
[0223]
[Table 11]

[0224]
Table 12 shows tensile properties at room temperature, V-notch Charpy impact absorption energy at 20 ° C., 650 ° C., 10⁵h shows the creep rupture strength and weld crack test results.
[0225]
The creep rupture strength and impact absorption energy of the material of the present invention to which appropriate amounts of B, Mo and W are added are the properties required for a high-temperature high-pressure turbine casing (625 ° C., 10⁵h Strength ≧ 8kgf / mm², 20 ° C. shock absorption energy ≧ 1 kgf−m). Especially 9kgf / mm²The above high values are shown. Moreover, no weld cracking is observed in the material of the present invention, and the weldability is good. As a result of investigating the relationship between the amount of B and weld cracks, when the amount of B exceeded 0.0035%, weld cracks occurred. No. One was a little worried about cracking. Looking at the influence of Mo on mechanical properties, those with a large Mo content of 1.18% had high creep rupture strength, but had low impact values and could not satisfy the required toughness. On the other hand, Mo 0.11% had high toughness but low creep rupture strength and could not satisfy the required strength.
[0226]
As a result of investigating the influence of W on the mechanical properties, creep rupture strength is remarkably increased when the W amount is 1.1% or more, but conversely, when the W amount is 2% or more, the room temperature impact absorption energy is lowered. . In particular, by adjusting the Ni / W ratio to 0.25 to 0.75, the temperature is 621 ° C. and the pressure is 250 kgf / cm.²625 ° C., 10 required for the high-pressure and medium-pressure inner casings of the above high-temperature and high-pressure turbines, and the main steam stop valve and the control valve casing.⁵h Creep rupture strength 9 kgf / mm²As described above, a heat-resistant cast steel casing material having a room temperature impact absorption energy of 1 kgf-m or more is obtained. In particular, by adjusting the W amount to 1.2 to 2% and the Ni / W ratio to 0.25 to 0.75,⁵h Creep rupture strength 10 kgf / mm²As described above, an excellent heat-resistant cast steel casing material having a room temperature impact absorption energy of 2 kgf-m or more can be obtained.
[0227]
[Table 12]

[0228]
The amount of W is remarkably strengthened by setting it to 1.0% or more, and especially at 1.5% or more, 8.0 kgf / mm.²The above values are obtained. No. of the present invention. 7 satisfied the required strength sufficiently at 640 ° C. or lower.
[0229]
An alloy raw material having the target composition of the heat-resistant cast steel of the present invention was melted by 1 ton in an electric furnace, and after ladle refining, it was cast into a sand mold and the inner casing of the high and medium pressure part described in Example 3 was obtained. The casing was annealed at 1050 ° C. × 8 h for furnace cooling, tempered at 1050 ° C. × 8 h for blast cooling, and tempered twice at 730 ° C. × 8 h for furnace cooling. As a result of cutting investigation of this prototype casing having a fully tempered martensite structure, the characteristics required for a 250 atm, 625 ° C high temperature and high pressure turbine casing (625 ° C, 10⁵h Strength ≧ 9 kgf / mm², 20 ° C. impact absorption energy ≧ 1 kgf−m) and sufficiently weldable.
[0230]
(Example 5)
In the present embodiment, the steam temperature of the high-pressure steam turbine and the intermediate-pressure steam turbine or the high-medium-pressure steam turbine is changed to 649 ° C. instead of 625 ° C., and the structure and the size are almost the same as those in the second or third embodiment. It is obtained by. Here, what is different from the first embodiment is the rotor shaft, the first stage moving blades and the first stage stationary blades, and the inner casing of the high pressure, intermediate pressure, or high intermediate pressure steam turbine that are in direct contact with this temperature. Among these materials excluding the inner casing, among the materials shown in Table 7 above, the B amount is increased to 0.01 to 0.03% and the Co amount is increased to 5 to 7%. By increasing the amount of W to 2 to 3% and adding 3% of Co, the required strength is satisfied and there is a great merit that the conventional design can be used. That is, in the present embodiment, the conventional design concept can be used as it is in that all the structural materials exposed to high temperature are made of ferritic steel. Since the steam inlet temperatures of the second stage moving blade and stationary blade are about 610 ° C., it is preferable to use the material used in the first stage of Example 1.
[0231]
Further, the steam temperature of the low-pressure steam turbine is about 405 ° C., which is slightly higher than about 380 ° C. of Example 2 or 3, but the rotor shaft itself has a sufficiently high strength, so Clean material is used.
[0232]
Furthermore, the cross compound type in the present embodiment is a tandem type that is directly connected to the cross compound type, and can be carried out at a rotational speed of 3600 rpm.
[0233]
【The invention's effect】
According to the invention,Rotor shaft for steam turbine having journal portion having high weldability in rotor shaft made of martensitic steel with higher alloyingIs obtained.
[Brief description of the drawings]
FIG. 1 is a cross-sectional view of a high-pressure and intermediate-pressure steam turbine according to the present invention.
FIG. 2 is a sectional structural view of a low-pressure steam turbine according to the present invention.
FIG. 3 is a perspective view of a turbine rotor blade according to the present invention.
FIG. 4 is a configuration diagram of a coal combustion power plant according to the present invention.
FIG. 5 is a cross-sectional view of a rotor shaft for a high pressure steam turbine according to the present invention.
FIG. 6 is a cross-sectional view of an intermediate pressure steam turbine rotor shaft according to the present invention.
FIG. 7 is a cross-sectional view of a high intermediate pressure steam turbine according to the present invention.
FIG. 8 is a cross-sectional view of a high intermediate pressure steam turbine rotor shaft according to the present invention.
FIG. 9 is a cross-sectional view of a low-pressure steam turbine according to the present invention.
FIG. 10 is a sectional view of a rotor shaft for a low pressure steam turbine according to the present invention.
[Explanation of symbols]
DESCRIPTION OF SYMBOLS 1 ... 1st bearing, 2 ... 2nd bearing, 3 ... 3rd bearing, 4 ... 4th bearing, 5 ... Thrust bearing, 10 ... 1st shaft packing, 11 ... 2nd shaft packing, 12 ... 3rd shaft packing, DESCRIPTION OF SYMBOLS 13 ... 4th shaft packing, 14 ... High pressure partition plate, 15 ... Medium pressure partition plate, 16 ... High pressure moving blade, 17 ... Medium pressure moving blade, 18 ... High pressure internal casing, 19 ... High pressure external casing, 20 ... Medium Pressure inner first casing, 21 ... Medium pressure inner second casing, 22 ... Medium pressure outer casing, 23 ... High pressure axle, 24 ... Medium pressure axle, 25 ... Flange, elbow, 26 ... Front bearing box, 27 ... Journal part, 28 ... main steam inlet, 29 ... reheat steam inlet, 30 ... high pressure steam exhaust port, 31 ... cylinder connecting pipe, 38 ... nozzle box (high pressure first stage), 39 ... thrust bearing wear shut-off device, 40 ... Warm-up steam inlet, 43 ... bearing, 51 ... boiler, 52 ... high pressure turbine, 53 ... medium Turbine, 54, 55 ... Low pressure turbine, 56 ... Condenser, 57 ... Condensate pump, 58 ... Low pressure feed water heater system, 59 ... Deaerator, 60 ... Boost pump, 61 ... Feed water pump, 63 ... High pressure feed water heating System: 64 ... economizer, 65 ... evaporator, 66 ... superheater, 67 ... air heater, 68 ... generator.

Claims (5)

A rotor shaft for a steam turbine made of martensitic steel in which a rotor journal part and a rotor body part are integrally formed,
The C0.06～0.14% rotor journal portion in a weight ratio, Si0.5% or less, Mn2% or less, Cr7~ 13%, Ni0.2~2%, V0.05~ 0. 35%, Nb0. 01~0.20%, N0.005~ 0. 05% , Mo1% or less, W3% or less, B no additives or 0.0030% or less and viewed including the Co5% or less, from martensitic steel of the rotor barrel Made of martensitic steel with a high weldability alloy composition ,
C0.06～0.14% rotor body portion in a weight ratio, Si 0. 15% or less, Mn 0. 03~1. 5% , Cr 9~13%, Ni 0. 05~1. 0%, V0 .05~0.3%, Nb0.01~0.20%, N0.005~ 0 . 05%, Mo 0. 05~0. 5%, W 1. 0~3. 5%, B 0. 001 comprises 0.03% and Co 10% or less, the steam turbine rotor shaft, characterized in <br/> be composed of martensite steel having a high alloy composition-temperature strength than the martensite steel of the rotor journal. The rotor shaft according to claim 1 , wherein a low alloy steel having good bearing characteristics is welded to the surface of the journal portion. In the manufacturing method of the rotor shaft of Claim 1 , 2 or more types of consumable electrodes of the martensitic steel with good weldability and the martensitic steel with high high-temperature strength are prepared, and the electrodes are obtained by remelting the electroslag. A method of manufacturing a rotor shaft for a steam turbine, wherein a journal part and a body part are manufactured integrally. 4. The steam turbine rotor according to claim 3 , wherein a consumable electrode corresponding to a journal portion at both ends and the barrel portion at the center is prepared, and the consumable electrode is electroslag-melted. Shaft manufacturing method. 4. The method of manufacturing a rotor shaft according to claim 3 , wherein the consumable electrode of martensitic steel with good weldability is first melted, then the consumable electrode of martensitic steel with high high-temperature strength is melted, and further the weldability is good A method for manufacturing a rotor shaft for a steam turbine, wherein a consumable electrode made of martensitic steel is melted and manufactured integrally.

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